CONTENTS

SCIENTIFIC AND TECHNICALKuchuk-Yatsenko S.I., Shvets Yu.V. and Shvets V.I.Influence of non-metallic inclusions in pipe steels ofstrength class X65—X80 on values of impact toughnessof flash-butt welded joints ....................................................... 2

Osin V.V. Tribotechnical properties of deposited metalof 50Kh9S3G type with increased sulphur content ................... 8

Demidenko L.Yu., Onatskaya N.A. and Polovinka V.D.Effect of temperature of thermomechanical treatmenton quality of dissimilar metal joints .......................................... 11

Borisov Yu.S., Vojnarovich S.G., Kislitsa A.N. andKalyuzhny S.N. Investigation of spraying spot andmetallization pattern under conditions of microplasmaspraying of coatings of titanium dioxide ................................... 15

Korzhik V.N., Borisova A.L., Popov V.V., KolomytsevM.V., Chajka A.A., Tkachuk V.I. and Vigilyanskaya N.V.Cermet coatings of chromium carbide—nichromesystem produced by supersonic plasma gas airspraying ................................................................................. 19

INDUSTRIALZyakhor I.V., Zavertanny M.S. and Chernobaj S.V.Linear friction welding of metallic materials (Review) ................ 25

Chvertko P.N., Semyonov L.A. and Gushchin K.V.Flash-butt welding of thin-walled profiles ofheat-hardened aluminium alloys .............................................. 32

Zhudra A.P., Krivchikov S.Yu. and Dzykovich V.I.Application of complex-alloyed powders produced bythermocentrifugal sputtering in flux-cored wires ....................... 36

Pereplyotchikov E.F., Ryabtsev I.A., Lankin Yu.N.,Semikin V.F. and Osechkov P.P. Modernization ofcontrol system of A1756 machine for plasma-powdersurfacing ................................................................................ 41

INFORMATIONTraining of specialists-welders of Kazakhstan in Ukraine ........... 45

Index of articles for TPWJ’2014, Nos. 1—12 .............................. 46

List of authors ........................................................................ 50

English translation of the monthly «Avtomaticheskaya Svarka» (Automatic Welding) journal published in Russian since 1948

Editor-in-Chief B.E.Paton

EDITORIAL BOARDYu.S. Borisov,

B.V. Khitrovskaya (exec. secretary),V.F. Khorunov, V.V. Knysh, I.V. Krivtsun,S.I. Kuchuk-Yatsenko (vice-chief editor),

Yu.N. Lankin, V.N. Lipodaev (vice-chief editor),L.M. Lobanov, A.A. Mazur,

I.K. Pokhodnya, V.D. Poznyakov, I.A. Ryabtsev,K.A. Yushchenko,

A.T. Zelnichenko (exec. director)(Editorial Board Includes PWI Scientists)

INTERNATIONAL EDITORIALCOUNCIL

N.P. AlyoshinN.E. Bauman MSTU, Moscow, Russia

V.G. FartushnyWelding Society of Ukraine, Kiev, Ukraine

Guan QiaoBeijing Aeronautical Institute, China

V.I. LysakVolgograd State Technical University, Russia

B.E. PatonPWI, Kiev, Ukraine

Ya. PilarczykWeiding Institute, Gliwice, Poland

U. ReisgenWelding and Joining Institute, Aachen, Germany

O.I. SteklovWelding Society, Moscow, Russia

G.A. TurichinSt.-Petersburg State Polytechn. Univ., Russia

M. ZinigradCollege of Judea & Samaria, Ariel, Israel

A.S. ZubchenkoOKB «Gidropress», Podolsk, Russia

FoundersE.O. Paton Electric Welding Institute

of the NAS of Ukraine,International Association «Welding»

PublisherInternational Association «Welding»

TranslatorsA.A. Fomin, O.S. Kurochko,

I.N. KutianovaEditor

N.A. DmitrievaElectron galley

D.I. Sereda, T.Yu. Snegiryova

AddressE.O. Paton Electric Welding Institute,International Association «Welding»

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December 2014No.12

Published since 2000

E.O. Paton Electric Welding Institute of the National Academy of Sciences of Ukraine

International Scientific-Technical and Production Journal

© PWI, International Association «Welding», 2014

INFLUENCE OF NON-METALLIC INCLUSIONSIN PIPE STEELS OF STRENGTH CLASS X65—X80

ON VALUES OF IMPACT TOUGHNESSOF FLASH-BUTT WELDED JOINTS

S.I. KUCHUK-YATSENKO, Yu.V. SHVETS and V.I. SHVETSE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

The comprehensive mechanical tests of high-strength steel pipe joints produced using flash-butt weldingwere carried out. The quality of joints completely meets the requirements of the international standardAPI STANDARD 1104. During tests on impact bending the single dropouts of KCV—40 values lower thanthe level required by standards DNV-OS-F101 and STO Gazprom 2-3.7-380—2009 were observed. Thedecrease in KCV—40 was predetermined by formation of zones of 1—2 mm2 area in the plane of a jointdiffered by structural heterogeneity, which is characteristic also for base metal and revealed in a sharplydistinct anisotropy of its ductile properties. It is shown that the areas with structural heterogeneity do notform colonies, therefore they do not decrease the general serviceability of the welded joint. The sizes ofsuch areas do not exceed the admissible values accepted by the standards for defects of welds made by arcwelding. The presence of single local dropouts of KCV—40 values observed during tests of welded jointsproduced using FBW should not be considered as a sign of rejection. 8 Ref., 1 Table, 11 Figures.

Keywo r d s : flash-butt welding, high-strength steel,impact toughness of welded joints, heat treatment ofwelds

The many-year works in flash-butt welding(FBW) of large-diameters pipes [1] at the E.O.Paton Electric Welding Institute have found thefurther development over the recent years. Thetechnology of FBW of thick-walled pipes of1219—1420 mm diameter of high strength steelsused in construction of modern pipelines of highefficiency was developed. The technology ofwelding and heat treatment provides the jointswith high values of mechanical properties, whichmeet the requirements of international standards[2—4]. Basing on this technology CJSC «Pskov-elektrosvar» (Russia) jointly with the E.O. Pa-ton Electric Welding Institute developed a newgeneration of equipment for FBW of pipes of1219—1420 mm diameter. At the present time thecomprehensive tests of this equipment and weld-

ing technology and its adaptation to standardrequirements specified for the quality of weldedjoints during construction of off-shore and on-shore pipelines including under conditions of theExtreme North are carried out.

For such pipelines the International standards[3, 4] specified the standard values of impacttoughness of welded joints of circumferentialwelds of not less than 37.5 J/cm2 at the testtemperature, which should be 20 °C lower thanthe design temperature of pipeline service.

The aim of the present article is determinationof factors, which influence the value and stabilityof values of impact toughness of high-strengthsteel pipe joints produced by FBW.

The development of welding technology wascarried out on the sectors of pipes of 1219 mmdiameter with the wall thickness of 27 mm, whichare manufactured of sheet steel 10G2FB of X70strength class, produced by controllable rolling

© S.I. KUCHUK-YATSENKO, Yu.V. SHVETS and V.I. SHVETS, 2014

Mechanical properties of metal of pipes and welded joints

Investigation areaYield strength,

MPaTensile

strength, MPa

Impact toughness, J/cm2, at T, °C

+20 —20 —30 —40

Base metal 484.4—493.5490.0

546.7—556.8553.0

334.7—336.6335.8

— — 333.0—336.6334.9

Welded joint after heat treatment — 550.6—561.4554.6

147.9—219.5173.2

86.8—171.1137.9

84.5—167.5115.5

19.1—129.198.6

2 12/2014

with thermal mechanical strengthening, of thefollowing chemical composition, wt.%: 0.06 C;0.21 Si; 1.42 Mn; 0.12 Ni; 0.07 Mo; 0.04 V;0.04 Al; 0.02 Ti; 0.05 Cr; 0.02 Nb; 0.004 S;0.012 P. The mechanical properties of base metaland also of welded joints after heat treatmentare given in the Table. The modes of weldingand heat treatment are presented in [5, 6]. Allthe welded joints were subjected to non-destruc-tive testing (radiographic, ultrasonic) of a highresolution capacity.

The mechanical tests of welded butts werecarried out at the PWI laboratory according tothe requirements stated in work [3]. This labo-ratory is certified according to the Internationalstandards. Metallographic examinations werecarried out in the light microscope «Neophot-32»,analyses of fracture surfaces and microstructureof joints were made in the JEOL Auger-micro-probe JAMP 9500F (Japan) in the laboratory ofmetallographic investigations.

In the specimens, tested for static bendingafter heat treatment and presented in Figure 1,the cracks and fractures were not detected.

As is seen from the Table and Figures 1, 2 thevalues of impact toughness in heat-treatedwelded joints meet the requirements in the tem-perature range of +20— —30 °C [3, 4], specifiedfor the off-shore pipelines. With the reductionof testing temperature Ttest of specimens for bend-ing tests the values of KCV are reduced andscattering of their values increases (see Fi-

gure 2). Nevertheless at Ttest= —30 °C they remainat a sufficiently high level of 84.5—167.5 J/cm2

and at Ttest = —40 °C, which is envisaged forcontinental pipelines under the conditions of theExtreme North, scattering of values increases andsingle dropouts below the level of 37.5 J/cm2

are observed, regulated by the standard DNV-OS-F101. The number of such specimens testedat the mentioned temperature does not exceed10—15 %. Here, the average value of KCV re-mains sufficiently high – 98.6 J/cm3.

The testing of welded joints of all the speci-mens using non-destructive methods did not de-tect any imperfections of metal structure in theplane of the joint, which could be classified asdefective ones, even at setting up the devices toincreased sensibility.

Macro- and microstructure of the pipe joint,produced at the optimal mode with the furtherheat treatment, are given in Figure 3. The totalHAZ width after welding and heat treatmentamounts to about 60 mm. The base of microstruc-ture of HAZ metal, as well as base metal is ferrite.Along the joint line the microstructure is featuredby somewhat large size of ferrite grain (10—15 μm) as compared to HAZ microstructure andpresence of partially decayed residual austenitewith formation of grain bainite.

In the course of comparison of weld micro-structure of specimens, which showed minimum

Figure 3. Macro- (a) and microstructure (b – ×1000) of joints of pipe steel at the area of joint line

Figure 1. Specimens of welded joint tested on static bendingafter heat treatment

Figure 2. Temperature dependence of impact toughness ofwelded joints

12/2014 3

and maximum values of KCV—40 during tests, nodifferences were revealed.

The analysis of structure of fractures of weldedspecimens after tests on impact bending was per-formed. The investigations showed that fracturesof specimens are crystalline in all the cases (Fi-gure 4). The developed relief in some cases isconnected with presence of coarse non-metallicinclusions (NMI) in metal.

Microstructure of fracture surfaces of all thespecimens is morphologically the same (Fi-gure 5). The facets of chips with the structureelements typical of cleavage fracture, such assteps of river pattern and lugs, are combined withtear ridges, i.e. the elements of tough fracture.The presence of secondary cracks along the in-terfaces are characteristic.

The impact toughness of specimens with suchmicrostructure at almost absent NMI correspon-ded to standard requirements (see Figure 4, a).

In the specimens with a low impact toughnessthe cluster of NMI was revealed at the fracturesurface (see Figure 4, c). X-ray spectral mi-croanalysis of chemical composition allowed dis-tinguishing the following types. These are thecomplex oxides of silicon, calcium, aluminiumof up to 50 μm size (Figure 6, spectra 1, 2),

particles on the base of silicon (Figure 6, spectra3, 4). These inclusions are refractory products ofmetallurgical reactions, which transfer to theweld almost without suffering the changes underthe thermodeformational conditions of welding.Impurity and alloying elements such as sulfur,niobium and titanium also can be included totheir composition.

Another type of the observed inclusions, form-ing clusters, are silicates. The areas with the par-ticles of silicates, which could be observed, oc-cupied the area up to several square millimeters.The fused appearance and multitude of particlesgive grounds to suppose that they represent thefragments of a liquid film crushed in a butt (Fi-gure 7).

To evaluate the effect of NMI in steel on struc-ture and composition of inclusions in welds, themetallographic examinations of base metal werecarried out.

Figure 4. Macrostructure of fractures of welded specimens after impact bending tests: a – KCV—40 = 110; b – 29.1;c – 19.1 J/cm2

Figure 5. Microstructure of fracture surface of welded speci-mens after impact bending tests

Figure 6. Refractory NMI at the fracture surface of weldedjoint, and results of their X-ray spectral microanalysis (hereand after – at.%)

4 12/2014

Microstructure of base metal represents a fer-rite matrix with negligible amount of pearlitecolonies (Figure 8, a). The size of grains of ferriteamounts by estimation to the limits from severalto 10 μm (11—12 according to ASTM), here thecoarse grains of ferrite are elongated along thedirection of rolling. The characteristic is the con-siderable level of foliation of microstructure,which is observed due to stringer-type locationof products of the eutectoid transformation.

NMI which are mainly represented by complexoxides and sulfides of aluminium, silicon, cal-cium, manganese, iron (Figure 8, b), in micro-structure are distributed non-homogeneously andare often concentrated in the strips of rolling.

To determine the influence of structural het-erogeneity of the pipe base metal on impacttoughness the tests of specimens, cut out frombase metal at different position of notch inCharpy specimens relatively to the strips of roll-ing, were carried out. Thus, in case of notch onthe surface of pipe across the direction of rollingat Ttest = —40 °C, KCV = (333.0—336.2)/334.9,and with notch in the middle of the surface ofthe pipe end along the central line of rolling,KCV = (9.8—236.5)/53.5 J/cm2.

The impact toughness tests of specimens ofbase metal with notch along the central line ofrolling showed that the results do not meet therequirements [4, 5]. The characteristic fracturesof specimens are shown in Figure 9.

The values of tests of specimens with notchalong the strips of rolling differ sharply by theexpressed instability. Among nine specimenswith notch along the central line of rolling, sixones had the results lower than the required37.5 J/cm2 (at minimum value of 9.8 J/cm2).As is seen, in the limits of strip on single areasthe impact toughness is high. Local decrease ofplastic properties was caused by structure het-erogeneity of rolling strips. However, it can not

weaken the metal even in the limits of area com-measurable with the sheet thickness.

On the fracture of base metal, opened alongthe rolling strip, the numerous inclusions are ob-served (Figure 10). As to their nature they are

Figure 7. Manganese alumosilicates on the fracture surface of pipe joints

Figure 8. Microstructure of base metal (a), and NMI inbase metal (b)

12/2014 5

very close to those observed in the base metal.Their base is oxide structures of silicon, alu-minium, manganese, iron oxides and sulfides.

The most obviously that transfer of NMI fromthe rolling strips to the plane of joint occursthrough the melt, being formed on the surfaceof flashing during welding. The thickness of meltlayer during flashing of low-alloyed steels ischanged in the limits of 0.1—0.6 mm.

It is shown in work [7] that at the presenceof holes, filled with oxides of iron or any metalfillers, in the part edges being flashed the zone,enriched with filler, is formed in the melt at theplaces of their escape to the edges being flashed.During upsetting this zone of melt is deformedand its area is many times increased.

In the studied case of formation of structuralhomogeneity in the plane, the role of filler isplayed by the content of liquation strip in theplace of its escape to the flashing surface.

During evaluation of mechanical properties ofwelded joints of large-diameter pipes accordingto the standards [3, 4] the values of tests ofstandard specimens, cut out of sectors of 300 mmwidth symmetrically on four regions of pipe pe-rimeter, are considered. Here, the results of me-chanical tests and values of non-destructive test-ing are compared. Basing on the analysis of thesedata on each sector the conclusion about admis-sion of dropouts of values of mechanical proper-ties and detection of defects is made.

During tests of welded joints of pipes, pro-duced using FBW at the optimal modes, no drop-outs from standard requirements on tensilestrength and also during tests on static bendingwere revealed.

To evaluate the distribution of areas with lowproperties of KCV in the full-scale specimen thecoordinate binding of position of impact speci-mens in the investigated sector of pipe of 300 mmwidth was made. Figure 11 shows results ofKCV—40 along the perimeter and across the thick-ness of welded butt.

As is seen, the impact specimens with the val-ues lower than the required ones, do not formthe colonies, but distributed stochastically in thesection of welded joint. Consequently, weldedjoint will not have the weakened areas noticeablyextended from the point of view of values oftoughness.

Figure 9. Fracture surface of Charpy impact specimens with notch along the central rolling line: a – KCV = 9.84; b –22.2; c – 30.8 J/cm2

Figure 10. Macro- (a) and microstructure (b) of base metalfracture, opened along the rolling strip, and results of X-rayspectral microanalysis

6 12/2014

All the given data give grounds to supposethat the presence of areas of structural heteroge-neity in welded joints and, correspondingly, localdecrease in values of impact toughness at the testedzone of welded joint area should not be consideredas the rejection mark during evaluation of the qual-ity of joints produced using FBW.

Using standards [4, 8] for electric arc weldingthe sizes of admissible defects and their totalextension at any area of weld of 300 mm at non-destructive testing are determined. The maxi-mum total length of such areas is 50 mm, andwidth is up to 1.5 mm as-applied to the defectsof the «lack of penetration» type. Even if oneaccepts that the areas with structural heteroge-neity in FBW are close to lack of penetrationsin arc welding, then the mentioned dropouts ofvalues of impact toughness in the butt welds arein the admissible ranges according to the stand-ard. Actually, the mechanical properties of areaswith structural heterogeneity are much higherand do not contain discontinuities, where metal-lic bond is absent at all.

Therefore the presence of single dropouts ofvalues of impact toughness at low test tempera-tures should not be referred to rejection signs.

Conclusions

1. The comprehensive mechanical tests of high-strength steel pipe joints produced by FBW werecarried out. The quality of joints completelymeets the requirements of the Internationalstandard API STANDARD 1104. It was foundthat in impact bending tests the separate dropoutsof KCV—40 values lower than the level requiredby standards DNV-OS-F101 and STO Gazprom2-3.7-380—2009 were observed.

2. Using means of non-destructive testingmethods of welded joints, none of welding de-

fects was revealed in the joints with low valuesof KCV—40. It was established that decrease invalues of KCV—40 is predetermined by formationof zones of 1—2 mm2 area in the plane of jointcharacterized by structural heterogeneity, typicalalso of the base metal, which is revealed in a sharplydistinct anisotropy of its ductile properties.

3. It is shown that the areas with structuralheterogeneity do not form colonies but are dis-tributed stochastically in the section of weldedjoint, therefore, they do not decrease the generalserviceability of welded joint. The sizes of suchareas do not exceed the admissible values ac-cepted by the standards for defects of welds pro-duced in arc welding.

4. The presence of separate local dropouts ofKCV—40 values, observed during tests of weldedjoints produced using FBW, should not be con-sidered as a rejecting sign.

1. Kuchuk-Yatsenko, S.I. (1986) Resistance butt weld-ing of pipelines. Kiev: Naukova Dumka.

2. API STANDARD 1104: Welding of pipelines and re-lated facilities. 21st ed.

3. DNV-OS-F101: Offshore standard. Submarine pipe-lines systems. Jan. 2000.

4. STO Gazprom 2-3.7-380—2009: Instruction on tech-nology of welding of marine pipelines. Moscow.

5. Kuchuk-Yatsenko, S.I., Shvets, Yu.V., Zagadarchuk,V.F. et al. (2012) Flash-butt welding of thick-walledpipes from high-strength steels of K56 strength class.The Paton Welding J., 5, 2—7.

6. Kuchuk-Yatsenko, S.I., Shvets, Yu.V., Zagadarchuk,V.F. et al. (2013) Technology of heat treatment ofpipe joints from steel of K56 grade produced byflash-butt welding. Ibid., 2, 2—7.

7. Kuchuk-Yatsenko, S.I., Kazymov, B.I., Zagadarchuk,V.F. et al. (1984) Formation of «dead spots» in jointproduced by resistance welding. Avtomatich. Svarka,11, 23—26.

8. STO Gazprom 22.4-359—2009: Instruction on non-de-structive testing of welded joints in construction ofland and underwater gas pipelines from steels X-80,X-100. Moscow.

Received 29.10.2014

Figure 11. Test results of specimens along the perimeter and across the thickness of welded butt on impact bending atTtest = —40 °C

12/2014 7

TRIBOTECHNICAL PROPERTIES OF DEPOSITED METALOF 50Kh9S3G TYPE

WITH INCREASED SULPHUR CONTENT

V.V. OSINE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

Steels alloyed with sulphur (for instance, silchrome) are widely applied in manufacture of tools for coldrolling of metal. Their reconditioning repair requires materials ensuring deposited metal composition closeto that of baze metal. The work is a study of sulphur influence on tribotechnical properties of depositedmetal of 50Kh9S3G type. Sulphur content in the deposited metal was varied in the range of 0.02 to 1.70 %.Investigations revealed that sulphur forms complex sulphides, which prevent seizure of interacting metalsurfaces, thus increasing deposited sample wear resistance. To ensure optimum tribotechnical properties ofdeposited metal, volume content of complex sulphides of the main alloying elements should be within 1.5to 2.0 %, and their dimensions should be ≤0.02 mm. This condition is ensured at total content of sulphurof 0.5—0.8 % in the deposited metal. At lower sulphur content the quantity of forming sulphides is insufficientto make an essential influence on deposited metal properties. At higher sulphur content the forming coarsesulphide inclusions are less strongly retained by deposited metal matrix and during testing they crumbleand are removed from the wearing zone. This adversely affects wear resistance of deposited metal andcontacting part. 10 Ref., 2 Tables, 3 Figures.

Keywo r d s : deposition, deposited metal, sulphur al-loying, sulphides, tribotechnical properties

Steel 50Kh9S3G, known as «silchrome» trademark, is quite widely applied for manufacturingmetal cold working tools, in particular, for deepcold drawing dies, as well as heavy-duty partsof some friction pairs. These parts operation proc-ess is characterized by significant mechanicalloads and they fail most often as a result of seizurewear. For reconditioning cladding of cold draw-ing dies PWI developed flux-cored wire PP-Np-50Kh9S3G, providing deposited metal similar tobase metal as to its composition [1]. Increase ofseizure resistance of deposited metal of this typeremains to be an urgent problem.

From scientific-technical publications [2—8]and earlier wear resistance studies at differentkinds of wear of 23Kh5M3FS deposited metal[9] alloyed with sulphur, it is known that inclu-sions of complex sulphides prevent seizure wear,lower the friction coefficient, and improve thequality of deposited metal wearing surfaces. Anobjective was set to improve the tribotechnicalcharacteristics of deposited metal of 50Kh9S3Gtype due to sulphur alloying.

Considering the experience of earlier studies[9], sulphur content in deposited metal of thistype was varied within 0.02—1.30 % (Table 1).

Wear resistance and friction coefficient of de-posited metal were assessed using an all-purposecomponent of friction machine, designed for labo-ratory-experimental evaluation of tribotechnicalproperties of friction pairs at room and elevatedtemperatures [10]. Wetting was assessed by thepresence or absence of increase of counterbodyor sample mass. Testing was conducted by holeabrasion method by «shaft-plane» schematicwithout additional lubricant feeding into frictionzone. Samples for tribotechnical studies, cut outof deposited metal third-fourth layer, had thedimensions of 3 × 17 × 25 mm. Wearing surfacewas 3 × 25 mm. Counterbody in the form of aring of 40 mm diameter and 12 mm width wasmade from quenched steel 45 with hardnessHRC 42.

Table 1. Composition and hardness of metal deposited with testflux-cored wires

Flux-cored wire designationWeight fraction of elements, % HRC

hard-nessC Mn Si Cr S

PP-Np-50Kh9S3G-Op-1 0.55 0.52 2.65 9.23 0.02 56

PP-Np-50Kh9S3G-Op-2 0.50 0.81 2.80 9.04 0.27 58

PP-Np-50Kh9S3G-Op-3 0.64 0.55 2.65 9.25 0.70 58

PP-Np-50Kh9S3G-Op-4 0.62 0.52 2.80 8.65 1.05 55

PP-Np-50Kh9S3G-Op-5 0.67 0.58 2.65 8.80 1.30 56

© V.V. OSIN, 2014

8 12/2014

The following test mode was selected by theresults of pre-testing: sliding velocity of0.06 m/s, 30 N load, test duration of 60 minafter run-in and sliding distance of about 227 m.This mode ensured stabilization in time of tri-botechical characteristics of all the studied ma-terials. At testing sample wear was determinedby the volume of abrasion hole, and counterbodywear – by the difference in its mass before andafter testing.

Wearing characteristics of deposited samplesand counterbodies at metal-over-metal frictionat room temperature are shown in Figure 1. Asis seen from the given data, wear of depositedmetal of 50Kh9S3G type at metal-over-metal fric-tion at room temperature first decreases with in-crease of sulphur content to 0.7 %, and then, atincrease of its content to 1.3 %, wear is intensifiedagain. Counterbody wear is similar to wear ofdeposited metal samples. No seizure wear wasdetected as shown by test results. In sample—counterbody friction pair the best wear resistancewas provided at sulphur content within 0.5—0.8 %. At increase of sulphur content in the sam-ple up to 1.3 % counterbody wear was also some-what increased.

Sulphur influence on friction coefficient of50Kh9S3G deposited metal was studied (Fi-gure 2). Deposited metal friction coefficient de-creases up to sulphur content of 0.7 % and then,with increase of sulphur content to 1.3 %, itremains at approximately the same level. As inthe case of wear, sulphur content of 0.5 to 0.8 %should be regarded as the optimum one.

A similar dependence of sample wear on sul-phur content was observed also in [9]. Appar-ently, sample wear resistance and, consequentlycounterbody wear resistance, is associated withmorphology and quantity of sulphide inclusions.

Investigation of microstructure of 50Kh9S3Gdeposited metal with different sulphur contentwas performed. Microstructure of 50Kh9S3G de-posited metal without sulphur consists of marten-site and residual austenite with a small amountof nonmetallic inclusions – silicates and oxides(Figure 3, a).

With addition of 0.27 % S, sulphide and oxy-sulphide inclusions appear in deposited metal mi-crostructure (Figure 3, b). With further increaseof sulphur content up to 0.7 %, deposited metalmicrostructure is refined, quantity and size ofsulphide and oxysulphide inclusions increasing(Figure 3, c). In samples with sulphur contentof 1.3 %, a multitude of individual sulphides ofglobular and elongated shape, as well as finesulphide clusters are observed (Figure 3, d).Scarce very coarse sulphides and oxysulphidesare also revealed, whereas practically no silicatesare observed.

Sulphur content influence on nonmetallic in-clusion volume fraction in unetched polished sec-tions was studied in quantitative analyzer «Om-nimet» at 600 magnification when scanningthrough 100 fields of view. At 0.7 % S contentin the deposited metal, volume fraction of sul-

Figure 1. Wear of samples of 50Kh9S3G deposited metalwith different content of sulphur (a) and counterbodiestested in pair with them (b)

Figure 3. Microstructure (×1000) of 50Kh9S3G depositedmetal with different sulphur content, %: a – 0.02; b –0.27; c – 0.70; d – 1.30

Figure 2. Sulphur influence on friction coefficient of50Kh9S3G deposited metal

12/2014 9

phides in upper deposited layers reaches 1.99 %(by volume), and at 1.3 % S content it is ≤3.21 %.

Dimensions of sulphide inclusions are greaterthan those of silicate ones, and sulphide inclu-sions become coarser with increase of sulphurcontent. At sulphur content of 0.7 %, sulphideinclusions size is from 0.007 to 0.02 mm, and atsulphur content of 1.3 % it is from 0.025 to0.05 mm, and single inclusions of 0.075 mm sizewere also found.

Microhardness of nonmetallic inclusions wasdetermined in the LECO hardness meter M-400at 10 g load. Sulphides have essentially lowermicrohardness than silicates, average microhard-ness of sulphides is 2835 MPa, and that of sili-cates is 9130 MPa. Results of X-ray microprobeanalysis show that manganese and chromium ac-tively react with sulphur, making up almost halfof sulphide phase volume (Table 2). Fraction ofiron sulphides is also high, while silicon is mostlypresent in the composition of complex sulphides.

Comparing the results of studying the micro-structure of 50Kh9S3G deposited metal with theresults of its wear resistance studies, leads to theconclusion that in order to ensure optimum tri-botechnical properties of 50Kh9S3G depositedmetal, volume fraction of complex sulphides ofthe main alloying elements in its structure shouldbe within the ranges of 1.5 to 2.0 %, and theirdimensions should be ≤0.02 mm. In order to sat-isfy these conditions, it is necessary for sulphurcontent in the deposited metal to be equal to 0.5to 0.8 %.

At lower content of sulphur the quantity offorming sulphides is insufficient for them to makea significant influence on properties of50Kh9S3G deposited metal. At greater sulphurcontent, the forming coarse sulphide inclusionsare weakly retained by deposited metal matrix,and during testing they are crushed and removedfrom the wearing zone. This impairs the wearresistance of deposited metal and mating part.Moreover, sulphide crushing out can be promotedby the combination of low microhardness and

large volume fraction of sulphide inclusions, thatmay lead to inability to resist high contact pres-sures in the friction zone.

Conclusions

1. It is established that at sulphur alloying of50Kh9S3G deposited metal, sulphides of themain alloying elements prevent seizure of wear-ing surfaces under the conditions of dry metal-over-metal friction.

2. To ensure optimum tribotechnical proper-ties of 50Kh9S3G deposited metal, volume frac-tion of complex sulphides in its structure shouldbe equal to 1.5—2.0 %, and their dimensionsshould be ≤0.02 mm. To satisfy this condition,sulphur content in the deposited metal should bewithin 0.5—0.8 %. At lower sulphur content sul-phides cannot have any essential influence ondeposited metal tribotechnical properties. Athigher sulphur content the forming coarse sul-phides are weakly retained by deposited metalmatrix and crush out during wear and are re-moved from the wearing zone that impairs thewear resistance.

1. Ryabtsev, I.A., Kondratiev, I.A. (1999) Mechanisedelectric arc surfacing of metallurgical equipmentparts. Kiev: Ekotekhnologiya.

2. Fedorchenko, I.M., Pugina, L.I., Slys, I.G. et al.(1973) Bearing sulphurised metal-ceramic materialson the base of stainless steels. In: Friction and wearat high temperatures, 115—120. Moscow: Nauka.

3. Vinogradov, Yu.M. (1965) Sulphidizing, selenizingand tellurizing of steels, cast iron and alloys. Metal-lovedenie i Termich. Obrab. Metallov, 10, 36—41.

4. Kovalchenko, M.S., Sychev, V.V., Tkachenko, Yu.G.et al. (1973) Influence of temperature on frictioncharacteristics of some sulphides, selenides and tellu-rides of refractory metals. In: Friction and wear athigh temperatures, 133—138. Moscow: Nauka.

5. Artamonov, A.Ya., Barsegyan, Sh.E., Repkin, Yu.D.(1968) Examination of lubricating properties of mo-lybdenum and tungsten disulphides. Poroshk. Metal-lurgiya, 12, 53—58.

6. Samsonov, G.V., Barsegyan, Sh.E., Tkachenko,Yu.G. (1973) On mechanism of lubricating action ofsulphides and selenides of refractory metals. Fiz.-Khimich. Mekhanika Materialov, 9(1), 58—61.

7. Lunev, V.V., Averin, V.V. (1988) Sulphur and phos-phorus in steel. Moscow: Metallurgiya.

8. Osin, V.V., Ryabtsev, I.A. (2004) Effect of sulphuron properties of iron-base alloys and prospects of itsapplication in surfacing materials. The Paton Weld-ing J., 10, 18—21.

9. Osin, V.V., Ryabtsev, I.A., Kondratiev, I.A. (2006)Study of sulfur effect on properties of deposited met-al of Kh5MFS type. Ibid., 12, 12—15.

10. Ryabtsev, I.I., Chernyak, Ya.P., Osin, V.V. (2004)Block-module unit for testing of deposited metal.Svarshchik, 1, 18—20.

Received 20.10.2014

Table 2. Composition of sulphide phase of 50Kh9S3G depositedmetal with 0.7 % S

Depositedmetal type

Analyzedsection

Weight fraction of elements, %

S Si Mn Cr Fe

50Kh9S3G Sulphide 12.53 1.22 28.81 13.15 44.28

Matrix 0.03 3.40 0.59 9.27 86.26

10 12/2014

EFFECT OF TEMPERATURE OF THERMOMECHANICALTREATMENT

ON QUALITY OF DISSIMILAR METAL JOINTS

L.Yu. DEMIDENKO, N.A. ONATSKAYA and V.D. POLOVINKAInstitute of Pulse Processes and Technologies, NASU

43a Oktyabrsky Ave., 54018, Nikolaev, Ukraine. E-mail:iipt@iipt.com.ua

At present time, pressure welding methods acquire more and more distribution for quality joining ofdissimilar metals. The Institute of Pulse Processes and Technologies of the NAS of Ukraine has developeda new method of welding of materials in solid state using current pulses. In essence, it lies in the fact thatmetals being welded at ambient temperature are pressure contracted and specific amount of current pulsesof amplitude density not less than 1.2⋅109 A/m2 and duration not less than 2⋅10—4 s are passed throughdeformation zone for intensification of plastic deformation at stage of physical contact formation. At thatmutual deformation of surface projection provides for sufficiently tight contact of surfaces being weldedand allows for their isolating from ambient environment. A welded joint is formed at further heating bytype of diffusion one without application of special shield of the surfaces from oxidation. Aim of presentwork is an investigation of effect of temperature of thermomechanical treatment on quality of joints inwelding of dissimilar metals in solid state using current pulses of large density. Such research methods asmechanical shearing tests, optical metallography and micro X-ray spectral analysis were used at that. Itwas determined as the result of experiments that 800 °C temperature is the most favorable for providing aquality steel 20 plus copper M1 welded joint, since at that no discontinuities are present in joint zone andclearly expressed initial structure is preserved in metal itself. 9 Ref., 3 Tables, 5 Figures.

Keywo r d s : welded joints of dissimilar metals, tem-perature of thermomechanical treatment, current pulsesof large density, self-sealing of inter-contact zone, metal-lographic investigations, mechanical shearing tests

Application of metals and alloys dissimilar ontheir properties is typical for manufacture ofwelded joints in current commercial production[1, 2]. This allows for using specific propertiesof each of them in full extent, reducing consump-tion of expensive and rare metals, producing ofparts with high service properties and reducingtheir weight. In this connection, development ofnew, efficient and low-power consuming techno-logical processes for producing of quality weldedjoints from dissimilar metals is relevant.

The Institute of Pulse Processes and Tech-nologies (IPPT) of the NAS of Ukraine developeda new method of welding of metals in solid stateusing current pulses [3]. It is a variant of pressurewelding with preheating.

In essence, it lies in the fact that metals beingwelded at ambient temperature are pressure con-tracted and specific amount of current pulses ofamplitude density not less than 1.2⋅109 A/m2

and duration not less than 2⋅10—4 s are passedthrough deformation zone for intensification ofplastic deformation at stage of physical contactformation. At that mutual deformation of surfaceprojection provides for sufficiently tight contactof surfaces being welded and allows their isolating

from ambient environment. A welded joint isformed at further heating without application ofspecial shield of the surfaces from oxidation [4].

Steel 20 and copper M1 were selected forwelding of dissimilar metals using proposedmethod. First of all, they are widely used inmanufacture of technical parts in electrical tech-nologies, transport, machine building and, sec-ondly, they have significant differences in theirphysical-mechanical properties and crystal-lographic characteristics. Mutual solubility ofthese metals is relatively small and depends ontemperature [5], therefore it was relevant to in-vestigate their effect on quality of producedwelded joints.

Aim of present work is an investigation ofeffect of temperature of thermomechanical treat-ment (TMT) on quality of joints in welding ofdissimilar metals in solid state using currentpulses of large density.

Materials and experiment procedure. Rec-tangular plates of 170 mm length, 12 mm widthand 2.5 mm thickness were used as mock-up speci-mens for welding of dissimilar metals using cur-rent pulses of large density, at that zone of weld-ing with length l ≈ 22 mm were taken in themiddle of specimen (Figure 1).

Mechanical preparation of plate surfaces wascarried out immediately before welding. It liedin their grinding up to combs height (microir-regularities) H0 = 5—6 μm. Then, the plate sur-faces were degreased by acetone. Plastic defor-© L.Yu. DEMIDENKO, N.A. ONATSKAYA and V.D. POLOVINKA, 2014

12/2014 11

mation of microprojections over contact surfaceswas carried out by pressing of the specimens atambient (room) temperature in special device atcalculation nominal pressure equal 50 MPa. Pulsecurrent treatment of contact zone of pressed plateswas performed by means of current passing alongone steel plate. Parameters of pulse current wasthe same as in work [4], namely, amplitude ofcurrent density J = 1.85⋅109 A/m2, pulse duration(attenuating sine wave) τ ≈ 3⋅10—4 s, number ofpulses N = 90 and pulse sequence frequency 0.5 Hz.

Pulse-treated specimens together with devicewere set in a chamber electric furnace for furtherheating.

Welded joints in order to determine the effectof TMT temperature on their quality were heatedin the furnace to welding temperature, whichaccording to [6] was selected in 0.5—0.7 rangefrom melting temperature, i.e. at 750, 800 and850 °C and holding at indicated temperature dur-ing 20 min. At least three specimens are necessaryfor each temperature level.

Quality of the welded joints were estimatedusing the results of metallographic investigationsand mechanical tests of templates, cut out fromcontact zone of welded specimens (see Figure 1).At that, each template was cut along the crosssection. One part of the templates was used formetallographic investigations and another onefor mechanical tests.

The microsections for metallographic investi-gations were taken normal to contact plane.Preparation of sections to investigations lied in

mechanical grinding and further etching in 4 %alcoholic nital [7].

Welded joint quality was preliminary estimatedby condition of boundary along the joint line (bydiscontinuities presence). At that, estimation ofboundary condition was carried out along its width,which is determined using metallographic micro-scope «Microtech» of MMO-1600 model at 400-fold magnification. If discontinuities along thejoint line are present, their measurement is per-formed at 1000-fold magnification in 100 pointshomogeneously distributed along the joint line.Fraction (in percent expression) of that or othervalues of boundary width along the joint line wascalculated based on their results.

Distribution (diffusion) of elements was ex-amined using a method of micro X-ray spectrumanalysis in the cross section from base metal tojoint line as well as along the joint line fromcopper side.

Mechanical strength tests (shearing tests)were carried out in accordance to GOST 10885—85 for determination of stresses, promoting shear-ing deformation with fracture of welded jointalong welding line. For this purpose, the speci-mens for investigations were produced from partof the templates (Figure 2).

The tests were carried out at laboratory bench,designed for investigation of static characteristicsof materials and developed at IPPT using thedevice, scheme of which is given in Figure 3.This device allows for providing pure shearingstress in tested section of the specimen when ap-plying tensile stresses Fs to guide plates. Loadingof the specimens was carried out gradually withholding at each step during 3 min.

Strength of the joints was estimated on shear-ing stresses τsh = Fsh/Ssh, where Fsh is the shear-ing force of tested section, kN; Ssh is the area ofthe section to be sheared, mm2.

Results of experiments and their discussion.Earlier the investigations were carried on studyof change of surface projection under effect ofpulse current using the scheme of current passingindicated above with the help of the Philips scan-ning electron microscope SEM 515 (Holland).Their results showed that treatment of prelimi-nary compressed metal plates of steel 20 + copperM1 by current pulses of large density promotesfor significant deformation of their surface projec-tion, as a consequence of which area of physicalcontact covering 60 to 80 % of general contact areais formed [4]. Mutual deformation of surface pro-jection if applicable to investigated welding methodprovides for sufficiently tight contact of surfacesbeing welded, that allows for their isolating fromenvironment and creates the possibility of producingof welded joint without application of shielding me-dium at further heating. This fact was verified ex-perimentally in the same work.

Figure 1. Scheme of contraction of specimens for treatmentby current pulses

Figure 2. Scheme of specimen for shearing tests: 1 – copperplate section; 2 – steel plate

12 12/2014

Rise of temperature at stage of volumetric in-teraction is necessary for artificial increase ofductility of welded metals and acceleration ofdiffusion processes. Besides, the heat treatmentintensifies diffusible solution of hydrogen fromnear-surface metal in metal depth. As the resultoxidized layer is diffused and does not play a roleof barrier for emergence of dislocations in the jointzone. Rates of activation and setting of contactsurfaces are increased at that, and development ofinteraction of metals being joined takes place incontact plane with formation of strong metallicbonds as well as in volume of contact zone.

Table 1 gives the averaged results of measure-ment of shearing stresses of welded joints pro-duced at different temperatures.

Comparison of strength of welded joints andbase metal was carried out. For that, base metalspecimen (copper in given case), which is similarto welded one (see Figure 2), was subjected tothermo-deformation impact, simulating weldingconditions, and tested using the same procedure.Thus, relative shearing stress τ

_sh was determined

for all investigated welding temperatures asτ_sh(T) = τsh(T)/τ(T),

where τsh(T) is the shearing stress, received atwelding temperature, MPa; τ(T) is the shearingstrength of base metal after heat treatment onwelding mode, MPa.

Table 1 provides for calculation values of rela-tive value of shearing stress.

The results of mechanical tests determinedthat shearing strength of produced welded jointsis changed in a narrow range from 174.2 to180.7 MPa at variation of welding temperaturefrom 750 to 850 °C, that corresponds to copperstrength 0.95—0.98.

Figure 4 shows the microstructures of zonesof produced welded joints, and Table 2 gives theresults of change of boundary width and its por-tion along the joint line. It can be seen from theTable that rise of TMT results in reduction ofboundary width as well as increase of number ofpoints along the joint line, in which its minimumwidth was registered. Thus, 8 % of joint lineexamined length have defects of 0.5—1.0 μm sizeat 750 °C THT temperature. The rest 92 % haveboundary width less than 0.5 μm. Defects of poretype and discontinuities in the joint zone arealready absent at higher temperatures (800 and

850 °C), and joint zone itself represents a grainboundary oriented in plate of initial contact (Fi-gure 4, b, c).

However, comparison of microstructures ofproduced welded joints shows that rise of TMTtemperature to 850 °C results in grain growth insteel 20 in the joint zone as well as base metal(see Figure 4, c). This is undesirable moment,since it can promote reduction of welded jointstrength characteristics.

It can be concluded based on the results ofmetallographic examinations that TMT tempera-ture of 800 °C is the most preferable for providingof quality welded joint from steel 20 + copperM1. At that, no discontinuities are present in thejoint zone, clearly expressed initial structure ispreserved in the metal itself, and joint zone rep-resents a grain boundary oriented in plane ofinitial contact.

Micro X-ray spectral analysis of nature of dis-tribution of main elements in the cross sectionof central part of welded specimen, received afterTMT at 800 °C during 20 min, indicate that thecomposition is changed from 100 % Fe (or Cu)to 100 % Cu (or Fe) (Figure 5) in near-contactzone of the joint. At that, penetration depth ofcopper (concentration up to 1 %) in steel 20 met-al is around 15 μm from joint line, and that foriron in copper makes approximately 13 μm. Theresults obtained can be completely agreed withcurrent understanding that appearance of diffu-sion zones at joint boundary is typical for metalswith limited mutual solubility [8]. Their forma-tion is related with the fact that some amount ofcopper atoms penetrates (diffuse) in iron throughinterface as a result of specific solubility. Simi-larly, some amount of iron atoms diffuses in cop-per (Table 3).

The results, presented in Table 3 and Figure 5,show formation of solid solutions along the joint lineand in cross section from the side of steel 20 as well

Table 1. Results of mechanical shearing tests of templates ofsteel 20 + copper M1 welded joints

Heat treatmenttemperature T, °C

Shearing stress τsh,MPa

Relative shearingstress τ

_sh

750 176.0 ± 6.2 0.96

800 180.7 ± 3.5 0.98

850 174.2 ± 5.1 0.95 Figure 3. Scheme of device for shearing testing of specimens:1, 2 – upper and lower guide plates, respectively; 3 –pin; 4 – welded specimen; 5 – washer; 6 – nut

Table 2. Width of boundary and its percent content along jointline at different TMT temperatures

Heat treatmenttemperature T,

°C

Boundary width, μm

1.0 ~0.5 <0.5

750 7 1 92

800 — — 100

850 — — 100

12/2014 13

as copper M1. This is an indication of processesof interatomic interaction of contact surfaces. Atthat, no phase formations, capable to result inwelded joint embrittlement, were found.

Thus, welding method under investigation al-lows for producing of quality joints of steel 20 +copper M1 at lower parameters of thermal defor-mation than in traditional diffusion welding. Be-sides, it can be done without application of spe-cial shield of surfaces being welded from oxida-tion, for example, vacuum processing or usingcontrolled inert gas atmosphere. Isolation of weld-ing zone from ambient media in welding using cur-rent pulses is provided as a result of self-sealing of

inter-contact zone due to current-stimulated in-tensive plastic deformation of surface layers ofmetals being joined. Further heating of sealedspace promotes formation of autovacuum, pro-viding self-cleaning of surfaces from oxide films[9] and production of welded joint.

Conclusions

1. TMT temperature of 800 °C is the most favor-able for welding of steel 20 and copper M1 dis-similar materials in solid state using currentpulses, since no defects of discontinuity type arepresent in the joint zone, and clearly expressedinitial structure is preserved in the metal itself.

2. It is shown that zones of solid solutions areformed in cross section along the joint line fromthe side of steel 20 as well as copper M1. At that,no phase formations, capable to result in weldedjoint embrittlement, were found in these zones.

3. The results of mechanical tests indicate thatproduced welded joints have sufficiently highshearing strength comparable with base metalstrength.

1. Middeldorf, K., von Hofe, D. (2008) Trends in join-ing technology. The Paton Welding J., 11, 33—39.

2. Reisgen, U., Stein, L., Steiners, M. (2010) Stahl-Aluminium-Mischverbindungen: Schweissen oderLoeten? Die Kombination zweier etablierterFuegetechnologien macht Unmoegliches moeglich.Schweissen und Schneiden, 62(5), 278—284.

3. Polovinka, V.D., Yurchenko, E.S., Vovchenko, O.I.Method of diffusion welding of metals. Pat. 79181Ukraine. Int. Cl. B23K 31/02. Publ. 25.05.07.

4. Vovchenko, A.I., Demidenko, L.Yu., Onatskaya, N.A.(2011) Intensification of plastic deformation of metalsurfaces under action of current pulses in pressure weld-ing. Naukovi Notatky: Mizhvuz. Zbirnyk, 32, 63—68.

5. Kurt, A., Uygur, E., Mutlu, E. (2006) The effect ofallotropic transformation temperature in diffusion-welded low-carbon steel and copper. Metallofiz.Nov. Tekhnol., 28(1), 39—52.

6. Kazakov, N.F. (1976) Diffusion welding of materi-als. Moscow: Mashinostroenie.

7. Kovalenko, V.S. (1973) Metallografic reagents.Moscow: Metallurgiya.

8. Karakozov, E.S. (1976) Joining of metals in solidphase. Moscow: Metallurgiya.

9. Paton, B.E., Medovar, B.I., Saenko, V.Ya. (1980)Spontaneous cleaning of metals from oxide films.Doklady AN SSSR, 159(1), 1117—1118.

Received 05.05.2014

Figure 4. Microstructure (×400) of zone of steel 20 + copperM1 welded joints at T = 750 (a), 800 (b) and 850 (c) °C

Table 3. Distribution of elements along joint line from copperside (in %) at depth around 10 μm

PointNo.

Cu Fe Cr Mn Ni

1 99.60 0.38 — — 0.02

2 97.82 2.12 0.06 — —

3 96.84 2.74 — 0.11 0.32

4 96.82 3.14 — 0.03 0.02

5 97.22 2.45 — 0.13 0.20

6 96.40 3.27 0.12 0.02 0.22

7 96.85 3.09 — — 0.07

8 96.56 3.13 0.02 0.05 0.25

Figure 5. Nature of distribution of main elements in crosssection of central part of steel 20 + copper M1 welded jointafter TMT at T = 800 °C

14 12/2014

INVESTIGATION OF SPRAYING SPOTAND METALLIZATION PATTERN UNDER CONDITIONS

OF MICROPLASMA SPRAYING OF COATINGSOF TITANIUM DIOXIDE

Yu.S. BORISOV, S.G. VOJNAROVICH, A.N. KISLITSA and S.N. KALYUZHNYE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

Analyzed is the gained experience in application of gas-thermal technologies in producing of electrocon-ductive, dielectric and resistive coatings for machine building, instrument-making industry and otherindustry branches. It is shown that the most challenging method for formation of resistive coatings inmanufacture of heating elements is a method of the plasma-arc spraying. It was found that during theprocess of manufacture of the resistive heating elements of small sizes (for example, for radio electronics)by using the method of traditional plasma-arc spraying the losses of material being sprayed, caused by ageometric factor, are increased. In this connection, to increase the degree of applying of sprayed materials,the technology of microplasma spraying is the mostly challenging. The work is aimed at the study offormation of a spraying spot and a metallization pattern under the conditions of microplasma spraying ofthe titanium dioxide coating. It was found during investigations that the spraying spot of TiO2 has a formof ellipse with 6.0—9.2 mm sizes of axes, where the smaller axis is directed along the horizontal line, andthe larger one – along the vertical line. Ratio of axes is 1.01—1.47 and depends on the spraying modeparameters. Determined are the losses of material being sprayed due to a geometric factor, which were 53 %in spraying of path of 1 mm width and less than 1 % in spraying the path of 5 mm width. 19 Ref., 1 Table,5 Figures.

Keywo r d s : microplasma spraying, titanium dioxi-de, resistive heating element, metallization pattern

Today, a large experience has been gained in pro-ducing of different types of coatings, such aselectroconductive, dielectric and resistive onesusing the methods of gas thermal spraying(GTS), which are designed for different industrybranches (machine building, electrical engineer-ing, instrument-making etc.) [1—3].

The resistive coatings, produced by the GTSmethods, represent a great practical interest.Analysis of literature about carried out investiga-tions on application of GTS methods for formationof resistive coatings in electrical engineering con-firms many times the challenging future of thesetechnologies [4, 5]. With the development andimprovement of the GTS technology during therecent years a method of plasma-arc technologyfinds the more and more wide application in theproduction of the resistive coatings [6—9]. In par-ticular, this method was perfect in manufacture ofresistive heating elements (RHE) [10—12].

The RHE, manufactured by the method ofplasma-arc spraying, are characterized by a sig-nificant decrease in temperature in current-car-rying layers and increase in service life [13—15].

However, during the process of manufacture ofRHE of small sizes for radio electronics usingthe traditional plasma-arc spraying the addi-tional costs appear, caused by increase in lossesof material being sprayed, which are charac-terized as losses connected with a geometric fac-tor. These losses are predetermined by the factthat the spot diameter of spraying for the tradi-tional plasma-arc spraying is 20—25 mm, that ex-ceeds the path width (2—5 mm) of the RHE.

To increase the degree of application of thematerial being sprayed it is rational to apply thetechnology of a microplasma spraying (MPS)[16, 17]. This technology will allow producingcoatings from different types of materials, reduc-ing greatly the losses of material being sprayeddue to a small diameter of a spraying spot (3—5mm), having in this case the minimum thermaleffect on the substrate, thus allowing coatingsto be sprayed on thin-walled parts without theirdistortion.

In the present work the formation of the spray-ing spot and metallization pattern was studiedunder the conditions of microplasma depositionof titanium dioxide coatings, the losses of mate-rial being sprayed, caused by a geometric factor

© Yu.S. BORISOV, S.G. VOJNAROVICH, A.N. KISLITSA and S.N. KALYUZHNY, 2014

12/2014 15

depending on width of path being sprayed, weredetermined.

To analyze the material losses, connected witha geometric factor, a number of experiments wascarried out for determination of parameters ofthe metallization pattern, which describes thedistribution of the coating material mass in thespraying spot. The metallization patterns wereobtained during the process of spraying at a fixedplasmatron into a spot on flat specimens of 20 mmsize during 10 s, and then the measurements ofvertical (large) L and horizontal (small) axes lof spraying spot, as well as maximum height ofthe deposited hill h were made (Figure 1).

Using a digital camera the macrofilming ofprofiles of the metallization pattern was made inthe directions normal to its axes (Figure 2).

Then, the processing of image was made fordetermination of coordinates of the pattern pro-file. Using MathCad program the metallizationpattern was designed according to these coordi-nates and its describing function was determined,from which the pattern area for larger and smalleraxes was calculated. Having data about the sizesof metallization pattern, it is possible to deter-mine such a parameter as an angle of opening ofthe plasma jet:

β = 0.5 arctg L

2H, (1)

where L is the bead width; H is the distance ofspraying.

Losses of material, connected with geometricfactor, were determined as

Lg.f = (1 — Sx-x/Stot)⋅100, %, (2)

where Sx-x is the area of metallization patternconfined by part size; Stot is the total area ofmetallization pattern (Figure 3).

The spraying was performed in MPN-004 typeinstallation for MPS. «Metachim» powder of ti-tanium dioxide (TiO2) with different sizes of powderof 15—40 μm was used as a material for spraying.

The intervals of varying, values of mode pa-rameters being studied and results of the experi-ment for measurement of geometric parametersof the spraying spot are given in the Table.

It was found during investigations that thespraying spot of TiO2 powder has a form of anellipse with 6.0—9.2 mm size of axes, where thesmaller axis is directed along the horizontal lineand the larger one – along the vertical line.Ratio of axes is 1.01—1.47 and depends on pa-rameters of the spraying mode. Probably, thisshape of the spraying spot is caused by the factthat during the powder feeding from batcherMD-004 a gravity force, directed normal to thejet axis, has an effect on its particles. As thepowder particles are differed by the size and,

Figure 1. Scheme of metallization pattern investigation

Figure 3. Losses of material due to geometric factor

Figure 2. Metallization pattern profiles of TiO2 material:a – large axis; b – small axis

Figure 4. Transverse section of metallization patterns inaxes of spraying spot: a – section along large axis y == 2.21⋅e—0.56x

2

; b – section along small axis y == 2.21⋅e—0.6x

2

; 1 – diagram of Gauss distribution (calcu-lated); 2 – real metallization pattern profile

16 12/2014

consequently, the mass and aerodynamic resis-tance, they will penetrate into the jet for differ-ent depth under the gravity force action. Underthese conditions the produced spraying spot willhave a shape of ellipse, the larger axis of whichis located in the vertical plane, i.e. it coincideswith the direction of the gravity force action (seeFigure 1). The analysis of nature of the curvesshowed that the geometry of the metallizationpattern at MPS is described rather reliably bythe Gauss function [18]:

y = y0e—kx

2

, (3)

where y0 is the thickness of coating at the beadaxis; k is the coefficient of coating material con-centration in the spraying spot.

Using experimental data, obtained as a resultof measurements the metallization pattern pro-files (measurements were made in large L andsmall l axes, see Figure 1) were plotted, whichcoincided with Gauss curves, for different modesof spraying (Figure 4). Coefficient of correlationK was 0.9849—0.9992. Value of k varied in therange of 0.12—0.97.

As a result of the carried out calculations ac-cording to formula (1) it was found that the

angle of opening of the microplasma jet is withinthe ranges of 2—5.2°. The obtained results corre-late with values given in literature for laminarplasma jets [19].

Calculations of losses of material beingsprayed, caused by a geometric factor, dependingon size of the resistive path, were made for eachexperiment and are given in the form of histo-grams (Figure 5).

The analysis of histogram can state that theleast losses are provided in applying of the mode6 (see the Table). This mode is characterized bya minimum value of current and high consumptionof plasma gas, that reduces the temperature of theplasma jet, thus decreasing the spraying distanceto substrate without risk of its overheating. De-crease in the distance leads in turn to minimizingof the spraying spot. In this case the losses were53 % in spraying of path of 1 mm width and lessthan 1 % in spraying of path of 5 mm.

Conclusions

1. It was established as a result of study of theprocess of titanium dioxide coating formation un-der the MPS conditions that the geometric sizesof the spraying spot depend on the spraying proc-

Figure 5. Losses of spraying material connected with geometric factor depending on size of the resistive path: a – forlarge diagonal; b – for smaller diagonal

Parameters of metallization pattern depending on mode of TiO2 spraying

Modenumber

I, A Gpl, l/min H, mm Pp.c, g/minHeight of

metallizationpattern, mm

Large axis, mm Smaller axis, mm

1 45 120 200 1.8 2.21 9.2 7.5

2 45 120 100 0.6 1.04 6.3 4.7

3 45 60 200 0.6 0.53 8.2 7.4

4 45 60 100 1.8 2.31 7.2 5.6

5 35 20 200 0.6 0.35 8.8 6.2

6 35 120 100 1.8 2.27 8.6 6.6

7 35 60 200 1.8 0.81 7.2 7.0

8 35 60 100 0.6 1.34 7.6 5.4

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ess parameters. The spraying spot has a form ofellipse with 6.0—9.2 mm sizes of axes at 1.01—1.47ratio of axes, and metallization pattern is de-scribed by the Gauss distribution.

2. It was shown that it is possible to controlthe material losses, caused by geometric factor,by selection of MPS modes of resistive paths ofTiO2 material. Minimum losses are attained onthe conditions of plasmatron operation at themode 6, which is characterized by minimum cur-rent value and high consumption of plasma gasthat reduces the plasma jet temperature and,thus, decreasing the spraying distance to sub-strate without risk of its overheating with de-crease of the spraying spot. This allows providingthe minimum losses of material being sprayed topath from 1 up to 5 mm width that equal to 53 %in spraying of path of 1 mm width and less than1 % in spraying of path of 5 mm width.

1. Vashkevich, F.F., Spalnik, A.Ya., Pluzhko, I.A.(2009) Electrothermal insulation of inductors for in-ternal heating of tube shells. In: Building, materialsscience, machine building. Dnepropetrovsk: PGSA.

2. Borisov, Yu.S., Kislitsa, A.N. (2002) Microplasmaspraying using wire materials. The Paton WeldingJ., 3, 50—51.

3. Kovalenko, G.D., Zombzhitsky, A.P. (1980) Specif-ics of plasma spraying of electric heating coatingswith dielectric filler. Fizika i Khimiya Obrab. Mate-rialov, 4, 86—89.

4. Lyasnikov, V.N., Perov, V.V., Lavrova, V.N. (1977)Application of plasma-arc spraying of alundum inmanufacturing of cathode-heating assembly-unit. In:Electronic engineering, Series Microwave electronics,Issue 4, 85—87.

5. Baklanov, D.I., Belyajkov, I.N., Virnik, A.M. et al.Method of manufacturing of resistive heating ele-ment. Pat. 2066514 RF. Int. Cl. H 05 B 3/12. Publ.10.09.96.

6. Scheitz, S., Toma, L., Berger, L.-M. et al. (2011)Thermisch gespritzte keramische Schichtheizelemente.Thermal Spray Bull., 4, 88—92.

7. Lyasnikov, V.K., Bogatyrev, G.F. (1978) Plasmaspraying of powder materials on parts of electronic

devices. In: Review on electronic engineering, SeriesTechnology, industrial engineering and equipment,Issue 4, 62.

8. Robson, G.J. (1976) Applications of plasma sprayingin hard facing. In: Proc. of Public Session and Met-als Technology Conf. (Sydney), 6.5.1—6.5.12.

9. Hasui, A. (1975) Procedure of spraying. Moscow:Mashinostroenie.

10. Dostanko, A.P., Vityaz, P.A. (2001) Plasma proc-esses in production of products of electronic engi-neering, Vol. 3. Minsk: FU AIN FORM.

11. Baranovsky, N.D., Sharonov, E.A., Vannovsky, V.V.(1991) Electrical properties of plasma coatings ofplane heating elements. In: Proc. of Conf. on Ther-mal Spraying in Industry of USSR and Abroad(Leningrad, 27—29 May 1991), 60—61.

12. Griffen, L.A., Dyadechko, A.G. et al. (1990) Heat-ing elements for fittings, produced by thermal spray-ing method of powders. Poroshk. Metallurgiya, 5,102—104.

13. Ershov, A.A., Urbakh, E.K., Faleev, V.A. et al. (1995)Plasma deposition of resistive layers of strip electricalheater. In: Proc. of Conf. on Physics of Low Tempera-ture Plasma (Petrozavodsk), Pt 3, 409—411.

14. Anshakov, A.S., Kazanov, A.M., Urbakh, E.K.(1998) Creation of low temperature heater by methodof plasma spraying. Fizika i Khimiya Obrab. Materi-alov, 3, 56—61.

15. Smyth, R.T. (1973) Thermal spraying of fine pow-ders. In: Proc. of 7th Int. Metal Spraying Conf.(London, 1973), 89—95.

16. Borisov, Yu.S., Pereverzev, Yu.N., Bobrik, V.G. etal. (1999) Deposition of narrow-band coatings bymethod of microplasma spraying. Avtomatich. Svar-ka, 6, 53—55.

17. Kislitsa, A.N., Kuzmich-Yanchuk, E.K., Kislitsa,N.Yu. (2009) Producing of narrow bands by microp-lasma spraying method from Ni—Cr wire. In: Abstr. ofAll-Ukr. Sci.-Techn. Conf. of Young Scientists andSpecialists on Welding and Related Technologies(Kiev, 27—29 May, 2009), 94.

18. Vojnarovich, S.G. (2012) Investigation of shape andsize of spraying spot and pattern of metallizationunder conditions of microplasma spraying of coatingsfrom hydroxyapatite. Vestnik NUK, 3, 81—84.

19. Antsiferov, V.N., Bobrov, G.V., Druzhinin, L.K.(1987) Powder metallurgy and spraying of coatings.Ed. by B.S. Mitin. Moscow: Metallurgiya.

Received 21.10.2014

18 12/2014

CERMET COATINGSOF CHROMIUM CARBIDE—NICHROME SYSTEM

PRODUCED BY SUPERSONICPLASMA GAS AIR SPRAYING

V.N. KORZHIK, A.L. BORISOVA, V.V. POPOV, M.V. KOLOMYTSEV, A.A. CHAJKA,V.I. TKACHUK and N.V. VIGILYANSKAYAE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

It is a well-known fact that application of supersonic methods of thermal spray deposition results insignificant increase of service properties of parts. However, up to moment no trails were made on producingof carbide or cermet coatings with the help of supersonic methods of plasma spraying. Present work isdedicated to production of cermet coatings from chromium carbide—nichrome compositions using a methodof supersonic plasma gas air spraying (SPGAS). SPGAS technology has series of advantages in comparisonwith well-known technology of HVOF spraying. First of all, it concerns efficiency, economy and charac-teristics (temperature, speed) of gas jet. Structure and phase composition of produced coatings wereinvestigated. It is shown that spraying using supersonic gas jets promotes for increase of content of Cr7C3carbide and reduction of NiCr in the coatings as a result of oxidation and formation of oxides. Coatingfrom composite powder differs by higher level of density and homogeneity, has layered fine-laminatedstructure with inclusions of fine carbides, contains lower amount of oxide phase, but larger quantity ofchromium carbide Cr3C2 in comparison with coating from mechanical mixture Cr3C2 + NiCr. These coatingscan be recommended for application as wear-resistant at increased temperatures. 6 Ref., 6 Tables, 7 Figures.

Keywo r d s : supersonic plasma gas air spraying, cer-met coatings, chromium carbide, powders, microstruc-ture, phase composition, microhardness

Plasma coatings based on chromium carbides arecharacterized by combination of such propertiesas resistance to wear by abrasive particles andhard surfaces at high temperatures (540—840 °C),high resistance under conditions of fretting-cor-rosion and aggressive media (for example, in liq-uid sodium at 200—625 °C), heat resistance attemperatures to 980 °C, radiation resistance [1].

Cr3C2 mechanical mixtures (rarely its mixturewith Cr7C3 or Cr23C6 carbide) with such metalsas Ni, Co, Ni+Cr, NiCr at different combinationof component content (amount of metallic bindercan be from 6—8 to 45 wt.%) are used for spraydeposition of cermet coatings with chromium car-bide.

Main methods, which are used for spray depo-sition of coatings with chromium carbides, areplasma, detonation and HVOF spraying [1—3].

It is a well known fact that application ofsupersonic methods of thermal coating depositionresults in significant increase of their serviceproperties. However, up to moment no trails weremade on producing of carbide or cermet coatingswith the help of supersonic methods of plasmaspraying.

Present work is dedicated to production ofcermet coatings from chromium carbide—nichro-me compositions using a method of supersonicplasma gas air spraying (SPGAS), examinationof structure and phase composition of producedcoatings.

SPGAS technology has series of advantagesin comparison with well-known technology of

Table 1. Comparative characteristics of HVOF and SPGAS technologies

Technology

Gas consumption, m3/h Powder consumption, kg/h Coefficientof material

use, %

Jet characteristics

Propane O2 N2 AirMetallicalloys

Oxides Speed, m/sTemperature,

°C

HVOF 3—4 15—21 1 — Up to 23 — 40—75 1400—2700 2800

SPGAS 0.3—40 — — 10—40 Up to 50 Up to 20 60—80 3000 3200—6300

© V.N. KORZHIK, A.L. BORISOVA, V.V. POPOV, M.V. KOLOMYTSEV, A.A. CHAJKA, V.I. TKACHUK and N.V. VIGILYANSKAYA, 2014

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HVOF spraying. First of all, it concerns efficiency,economy (cost of gases) and characteristics of gasjet (temperature, speed) (Table 1).

Coating spray deposition in present work wascarried out on «Kiev-S» installation, developedby Gas Institute and PWI of the NAS of Ukraine.Air mixture with addition of propane (around4 vol.%) was used as plasma gas at followingtechnological parameters: I = 260 A, U = 360 V,

180 mm distance, air pressure made 4 atm andconsumption 20 m3/h.

Powders of two types, namely PP-53Cr3C225(Ni20Cr) and PP-53B Cr3C2(Ni20Cr)(Table 2) from «Bay State Surface Technologies,Inc.» (USA) were used as materials for coatingspray deposition.

Powders of particle size 15—44 μm were ap-plied for coating spray deposition.

Table 2. Composition of powders

Powder gradeContent of elements, wt.%

Cr Ni C Mn Fe Si Rest

РР-53 71.10 17.10 9.49 0.036 1.75 0.065 0.41

РР-53В Base 20 9.40 Not indicated

Figure 1. Appearance (a, c) and microstructure (×800) of particles (b, d) of PP-53 (a, c) and PP-53B (b, d) powders:1 – carbide; 2 – metallic; 3 – cermet particles

Table 3. Composition of particles of initial powders based on XSMA results

Powder particleContent of elements, wt.%

Cr Ni C O* Additives

Carbide (1 in Figure 1, a, c) 70.10—79.70 1.50—5.60 18.9—21 — Fe – 0.73

Metallic (2 in Figure 1, a, c) 52—71.10 24.23—24.39 — — W – 0.71; Fe – 1.45

Cermet (3 in Figure 1, b, d) 30.77—68.10 10.55—49.81 9.17—13.60 6.2—7.7 Mg – 0.48—1.24; Al – 0.6; Fe – 0.55;W – 1.13—1.45

*Found in separate particles.

20 12/2014

Examinations of initial powders showed thatthey differ by appearance, microstructure, micro-hardness and phase composition. Thus, PP-53powder consists of faceted particles (Figure 1 a,c) and being a mechanical mixture of carbide andmetallic particles, which vary vastly on micro-hardness (13300 ± 1500 and 1980 ± 600 MPa,respectively).

Powder PP-53B consists of spherical particles(Figure 1, b, d), and each particle, taking intoaccount structure and microhardness, includesmetallic as well as carbide phases (microhardnessof particles is varied in 5000—10500 MPa range).

Examination of composition of powder parti-cles using X-ray spectral microanalysis (XSMA)(Table 3) showed that nickel or carbon are vir-tually absent in PP-53 powder. According to

Figure 2. X-ray patterns of particles of Cr3C2 + NiCr powder: a – PP-53; b – PP-53B

Figure 3. Microstructure of coatings from PP-53 (a, c) and PP-53B (b, d) powder (a, b – Neophot microscope, ×400;c, d – CamScan electron-probe analysis)

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carbon amount in the carbide particles (19—21 wt.%), they are close to Cr3C2 (19.34 vol. %)carbide on composition. Moreover, Cr7C3 carbide(9.01 wt.% C) (Figure 2, a) was found in powderusing X-ray phase structural analysis (XPSA)method.

Particles of PP-53B powder contain all threeelements (Cr, Ni and C) as well as additives ofthese elements such as small amount of Mg, Al,

W, Fe and Ca (see Table 3). Some particles haveareas enriched by oxygen at the level of 6—7 wt.%.

At the same time, results of XPSA of PP-53Bparticles indicate that they do not differ fromPP-53 powder (Figure 2, b) on qualitative con-tent, and since each particle contains metallic aswell as carbide phases, the powder can be con-sidered as composite one, consisting from cermetparticles.

Metallographic examination of sprayed coat-ings determined no fundamental differences intheir structure depending on powder type (Fi-gure 3). The only thing that can be noted ishigher level of homogeneity of coatings from com-posite powder (PP-53B) in comparison with coat-ings from mechanical mixture (PP-53). Besides,some difference is observed in values of averagemicrohardness of coatings and nature of micro-hardness variation curves (Table 4, Figure 4).

Thus, two most probable values of microhard-ness can be marked on microhardness variationcurve of the coating, produced from PP-53 pow-der (mechanical mixture Cr3C2 and NiCr). Thiscan be an evidence of presence in the structure,as in initial powder, of zones with harder carbidephases (HV = 11500 MPa) and metal enrichedzones (HV = 9000 MPa). Microhardness vari-ation curve of the coating from composite powderPP-53B has only one probable value of micro-hardness 10500 MPa, i.e. average between twoindicated above.

XPSA of phase composition of sprayed coat-ings showed (Figure 5) that they contain chro-

Table 4. Characteristic of sprayed coatings

Powder gradeMicrohardness, MPa

Phase composition, wt.%Average Most probable

РР-53 8420 ± 1550 9000; 11500 Cr7C3 – 26; Cr3C2 – 11; NiCr – 48; NiCr2O4 – 15

РР-53В 9960 ± 1400 10500 Cr7C3 – 21; Cr3C2 – 15; NiCr – 53; NiCrO3 – 11

Figure 5. X-ray patterns of sprayed coatings: a – PP-53;b – PP-53B powder

Figure 4. Microhardness variation curves of plasma coatings from PP-53 (a) and PP-53B (b) powders

22 12/2014

mium carbide Cr3C2, Cr7C3, nichrome (Ni80Cr20)and oxides NiCr2O4 (for coating from mechanicalmixture) or NiCrO3 (for coating from compositepowder), the latter were not found in initial pow-ders. Presence of indicated phased in the sprayedcoatings conforms the results of local XSMA (Ta-ble 5) of separate structural elements of thesprayed coatings (Figure 6).

It can be noted when comparing intensity ofX-ray reflections of separate phases in initial

powders (see Figure 2) and in sprayed coatings(see Figure 5) that spraying promotes increaseof content of Cr7C3 carbide in the coating incomparison with Cr3C2 and NiCr content reducesas a result of oxidation and oxide formation. X-ray spectrum microanalysis of the sprayed coat-

Table 5. Content of elements, determined by XSMA method, in structural zones of coatings from PP-53 and PP-53B powders (acc. toFigure 6, a and b, respectively)

Spectrum tobe analyzed

Content of elements, wt.%Supposed phase

Cr Ni C O Fe

1 (а) 13.33 81.57 — — 4.30 NiCr

2 (а) 81.82 1.45 16.72 — — Cr7C3

3 (а) 89.40 — 10.60 — — Cr3C2

4 (а) 56.41 18.50 4.87 20.22 — Oxide

1 (b) 61.13 25.89 11.66 — — Zones of cermets (metal + carbide)

2 (b) 13.99 75.11 8.71 — 0.93

3 (b) 26.50 61.97 8.89 — 0.79

4 (b) 28.93 59.57 — 9.44 0.49 Oxide

Table 6. Results of XSMA of sprayed coating from PP-53B pow-der (acc. to Figure 7)

Spectrum to be

analyzed

Composition, wt.%

C O Cr Fe Ni W

2 1.91 — — 97.66 — —

3 9.10 6.10 52.61 2.62 29.58 —

4 9.72 5.32 49.80 3.93 30.39 0.83

5 8.56 6.55 55.21 2.23 26.75 0.68

6 8.38 7.15 55.76 2.59 25.38 0.73

Figure 6. Zones of coatings from PP-53 (a) and PP-53B (b) powders examined by XSMA method

Figure 7. Coating from PP-53B powder examined by XSMAmethod

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ings (Figure 7, Table 6) showed that averagedcomposition (spectrum 3) and content of separatezones on depth (spectra 4—6) are somewhat differfrom values of the initial powder (see Table 1).This takes place as a result of appearance of oxidephases in the coatings. At that, amount of oxygenin the case of PP-53B powder coating somewhatincreases in direction from base interface to outersurface of the coating, while carbon content isreduced to the contrary.

The similar dependence is observed for PP-53powder coating. This fact is related with rise oftemperature of coating being sprayed in propor-tion to build-up of coating layer. Comparison ofcharacteristics of sprayed coatings from PP-53and PP-53B powders allows for noting that coat-ing from composite powder differs by higher levelof density and homogeneity, has layered fine-la-mellar structure with inclusions of fine carbides,contains lower amount of oxide phase, but lagerquantity of chromium carbide Cr3C2 in compari-son with coating from mechanical mixtureCr3C2 + NiCr.

It is a well-known fact that refractory prop-erties, stability and hardness of chromium car-bides are reduced with decrease of carbon amountin them [4, 5]. Therefore, carbon loss is undesir-able in the processes of thermal spraying of cer-met chromium carbide based coatings. Reviewpaper [6] dedicated to effect of thermal sprayingmethods (HVOF, air plasma spraying (APS),vacuum plasma spraying (VPS) and detonation

spraying) on properties of the coatings from me-chanical mixture Cr3C2 + NiCr indicates thathardness, abrasive wear resistance and thermalresistance of the coatings depend on carbon lossin spraying. Thus, for example, VPS coatingshave higher hardness and abrasive wear resis-tance, but lower thermal resistance among oth-ers, and APS coatings with argon-helium plasmagas are superior on their properties to the coatingsusing argon-hydrogen mixture etc.

The results of present work allows for con-cluding that quality of Cr3C2 + NiCr compositioncoating can also be increased with the help ofcomposite powder of the same compositions in-stead of mechanical mixture.

1. Borisov, Yu.S., Kharlamov, Yu.A., Sidorenko, S.L.et al. (1987) Thermal coatings from powder materi-als: Refer. Book. Kiev: Naukova Dumka.

2. Guilemagy, M., Nutting, I., Llorca-Isern, N. (1996)Microstructural examination of HVOF chromium car-bide coatings for high temperature applications. J.Thermal Spray Techn., 5(4), 483—489.

3. Guilemagy, J.M., Espallargas, N., Suegama, P.H. etal. (2006) Comparative study of Cr3C2—NiCr coating.Corrosion Sci., 48, 2998—3013.

4. Kiffer, R., Benezovsky, F. (1968) Solid materials.Moscow: Metallurgiya.

5. (1986) Properties, production and application of re-fractory joints. Ed. by T.Ya. Kosolapova. Moscow:Metallurgiya.

6. Takenchi, J., Nakahira-Kobe, A., Takara, G. (1993)Barbezat-Wohlen Plasma-Technik. Cr3C2 + NiCr cer-met coatings. In: Proc. of Thermal Spray Conf.(Aachen, 1993), 1—14.

Received 17.09.2014

24 12/2014

LINEAR FRICTION WELDING OF METALLIC MATERIALS(Review)

I.V. ZYAKHOR, M.S. ZAVERTANNY and S.V. CHERNOBAJE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

Capabilities and prospects for application of linear friction welding for joining various metallic materials,namely steels, titanium alloys, high-temperature nickel alloys and composite materials are considered.Parameters of the mode of linear friction welding for different material combinations and their mechanicalproperties are given. Structure of joints in linear friction welding is similar to that for other friction weldingvariants. Width of characteristic joint sections (zones of dynamic recrystallization, thermomechanical andthermal impact) depends on mode parameters – welding time, axial force, vibration amplitude andfrequency. Modern tendencies in modelling of heat evolution and deformation in linear friction weldingare analyzed. Cost reduction and improvement of reliability of equipment for realization of this processremain an urgent problem. At present the field of industrial application of this process is limited to aerospacecompanies, where components of gas turbine engines from titanium alloys are joined, and in the futureapplication of high-temperature and nickel alloy and composite materials is possible. 35 Ref., 5 Figures.

Keywo r d s : linear friction welding, process stages,joint structure, modelling, titanium alloys, high-tempera-ture nickel alloys, composite materials

Over an almost 60 year period of research andindustrial application of friction welding (FW) inthe world, an extensive volume of information hasbeen accumulated on joining various structural ma-terials. About 20 variants of technological proc-esses using frictional heating to produce permanentjoints, surfacing, forming, welding up holes, pro-ducing riveted joints have been proposed and im-plemented in practice [1]. Rotational FW, whichwas invented back in 1956, became the most widelyapplied in different industries.

Method of vibrational FW was described prac-tically in time with the start of FW commercialapplication. This method uses relative recipro-cating displacement for surfacing and joiningparts with rectangular cross-section [2, 3]. In1969 the mechanism of reciprocating motion forlow-carbon steel welding has been patented [4].

Method of vibrofrictional welding, which waslater called linear friction welding (LFW), waswidely applied for joining thermoplastic products[5], whereas its application for welding metallicmaterials was restrained by complexity and con-siderable cost of development of reliable equip-ment. At the end of 1980s, the requirements ofleading enterprises in aerospace industry necessi-tated development of technologies and equipmentfor LFW of various alloys for creation of weldedrotors of aircraft gas turbine engines (GTE).

In 1990 the Welding Institute put into opera-tion the first specialized machine with electro-mechanical drive for LFW of metals. Successfulwelding of rectangular samples of 25 × 6 mmsection from carbon and stainless austeniticsteels, aluminium alloy 5154, titanium alloy Ti—6Al—4V [6—8] was the starting point for suchcompanies as Rolls Royce, MTU Aero Enginesand Pratt & Whitney showing active interest inLFW technologies. LinFic® concept of LFWequipment based on application of high-precisioncomputerized hydraulic force drives was devel-oped and implemented through combined effortsof a group of eight European companies. Sincethat time leading research centers and enterprisesof various countries started intensively workingon development of equipment, studying LFWprocess, and its industrial application.

The objective of this review is analysis of tech-nological capabilities and features of formation ofmetallic material joints in linear friction welding.

General characteristic of LFW process.LFW process schematic is shown in Figure 1,diagram of mode parameter variation in time –in Figure 2.

Mechanism of welded joint formation at LFWis similar to that at rotational FW [9]. Weldedjoint forms in the following sequence:

• initial rubbing-in of friction surfaces. Rela-tive displacement of surfaces results in breakingup of oxide and grease films on contacting mi-croprotrusions with opening of juvenile sections,with formation and immediate breaking up ofadhesion bridges;

• avalanche-like increase of the number of inter-acting microprotrusions, as well as actual contactarea and fast increase of temperature in the butt;© I.V. ZYAKHOR, M.S. ZAVERTANNY and S.V. CHERNOBAJ, 2014

12/2014 25

• upsetting start (driving plasticized metal outof the butt), increase of butt temperature up toa certain value;

• quasistationary (equilibrium) state of thefriction process. Heat evolution power, tempera-ture in the butt, and upsetting speed are at acertain set level;

• deceleration, i.e. controllable lowering ofthe speed of relative displacement to zero, duringwhich the welded joint is formed;

• forging. Formed joint is subjected to compres-sive deformation by an axial force, which, as arule, does not exceed the force applied at heating.

Unlike FW, at LFW the joint form is char-acterized by unusual flash (Figure 3). An essen-tially larger amount of flash forms on the facesin the direction of reciprocal displacement thatis due to displacement of plasticized metal pre-dominantly in the direction of vibrations.

LFW advantages. This method allows pro-ducing sound joints from various materials, such

as titanium and nickel alloys, different steels,aluminium and its alloys, composite materials(CM), etc. The main difference of LFW fromrotational FW is the possibility of solid-phasewelding of products with rectangular cross-sec-tion. LFW has several advantages, also inherentto other FW variants [6—9]:

• capability of joining parts from difficult-to-weld materials both in similar and in dissimilarcombinations;

• high efficiency;• high and stable quality of the joints;• hygienic value – absence of evolutions of

harmful gases, ultraviolet and electromagneticradiation;

• no need for application of filler materials,fluxes and shielding gases;

• joint forms in the solid phase, structuralchanges in the base metal proceed to a small depth;

• low energy consumption.Main disadvantages of LFW are the high cost

and complexity of equipment manufacturing, dueto the need for application of high-rigidity high-precision force drives and sophisticated comput-erized control systems [7—9].

LFW parameters and materials beingwelded. The main parameters of LWF processare frequency and amplitude of reciprocal vibra-tions, pressure at heating and forging, heatingand forging time, upsetting value at heating andtotal upsetting during welding (see Figure 2)[10]. Additional LFW parameters that can influ-ence joint formation are time of acceleration andstopping of vibrations.

Works [11, 12] present results of LFW studiesof Ti—6Al—4V titanium alloy (Ti-64). Frequencyof reciprocal vibrations in welding reached119 Hz for vibration amplitude of 0.92 to 3 mm.It is shown that to produce sound joints in LFWit is necessary to exceed a certain value of specificpower of heat evolution that is determined as

w = αfP2πA

, (1)

where α, f is the amplitude and frequency ofreciprocalting displacement, respectively; P isthe force; A is the blank cross-sectional area.

Figure 1. Schematic of LFW process

Figure 3. Welded joint of titanium alloy at LFW

Figure 2. Diagram of variation of LFW parameters in time:1 – amplitude; 2 – compressive force; 3 – upsetting

26 12/2014

Producing sound joints (without pores or ox-ide inclusions) is possible in the case of simulta-neous provision of certain critical conditions,namely, high enough values of vibration ampli-tude and frequency and welding pressure. By thedata of [12] an important parameter for soundweld formation is total upsetting of blanks, itsoptimum value being different for different struc-tural materials.

Work [13] gives the results of studies of LFWof aluminium, steel and high-temperature alloys,which were conducted on tee-joints of «pin-to-plate» type with pin rectangular cross-section of12 × 4 and 20 × 4 mm, respectively. It is shownthat three properties of metallic material, namelyhigh-temperature strength, heat conductivity andfriction coefficient, dependent on temperature, areimportant for LFW. Materials quickly loosingtheir strength to great depth at heating, in par-ticular, aluminium and aluminium alloy ofAlMgSi1 system, are only partially suitable forLFW. Their stable heating requires ensuring con-siderable amplitudes and frequencies of reciprocat-ing displacements at considerable welding force.

Study [14] gives a comparative characteristicof LFW and inertia FW. Investigations were con-ducted on samples from high-temperature alu-minium alloy Al—11.7 % Fe—1.2 % V—2.4 % Si,produced by powder metallurgy techniques, aswell as samples of this alloy with 2024-T351 alu-minium alloy. It is shown that LFW allows re-ducing the intensity of thermodeformationalwelding cycle compared to inertia FW. Jointstrength is about 81 % of that of joints producedby intertia FW that is related to large coarseningof the grain and lower hardness in the joint zone.LFW mode for rods of 25 mm diameter is asfollows: 50 Hz frequency, 2 mm amplitude, 4 mmupsetting, 30 and 50 kN heating and forgingforces, respectively.

In [6, 10, 13, 15] LFW capabilities at joiningvarious steels (carbon, stainless, high-strength)were studied. Welding mode parameters werevaried within the following ranges: 25—50 Hzvibration frequency, 1—3 mm amplitude, 40—240 MPa heating force, and 1—3 mm upsetting.Authors of [6, 13] note that LFW of carbon steelruns into the following problems: welded jointshave undercuts on both sides of the joint zone.This is attributable to low linear speed(0.5 m/s), whereas it should be not less than1 m/s for steels to ensure uniform heating andstable deformation of metal.

Tensile testing of welded samples from carbonsteel [15] showed ultimate strength of 539—592 MPa with fracture through base metal(BM). Increase of metal microhardness in the

joint zone by more than 2 times compared to BMis observed. At deviation of LFW parametersfrom optimum values towards their lowering,welded joints fail through the butt with lowstrength values.

Sound joints from stainless steel AISI 3L61[16] were produced at more than 160 MPa weld-ing pressure and not less than 40 Hz vibrationfrequency. Mechanical testing showed ultimatestrength of 549—610 MPa, relative elongation of40—49 % with fracture running through BM. Fail-ure of samples produced at less than 80 MPawelding pressures or at 25 Hz vibration frequencyran through the weld. It is established that maxi-mum upsetting speed, which influences theamount of δ-ferrite in the joint zone, is achievedat high welding pressures and values of amplitudeof 1.5 mm and frequency of 30 Hz.

The greatest number of publications is devotedto studying LFW of various titanium alloys [6,17—22] that is due to LFW practical applicationin manufacture of welded rotors of aircraft GTEcompressors («blisks»). Publication authorsstudied the following ranges of LFW parametersvariation: 0.5—3.0 mm amplitude, 15—150 Hz fre-quency, 50—190 MPa pressure at heating and320 MPa forging pressure.

In [6, 17] welded joints of Ti—6Al—4V tita-nium alloy produced by LFW were studied. Atrupture testing of welded samples of 25 × 6 mmsection [6] strength was 835 MPa with fracturerunning through BM. At bend testing fractureran in the joint zone at bend angle of 50°. Me-chanical testing of welded joints of 35 × 26 ×× 13 mm model samples produced with upsettinglimitation in welding of 1.75 mm, showed thefollowing results [17]: ultimate strength of 916to 1004 MPa, relative elongation of 13.7—15.2 %,yield point of 841—968 MPa. Microhardness in-crease in the joint zone up to HV 350 (comparedto HV 300 in BM) was found at ±0.6—1.0 mmdistance from weld axis, depending on weldingpressure.

Work [18] is a study of the possibility of LFWapplication for new titanium pseudo-β-alloy Ti-5553 (Ti—5Al—5V—5Mo—3Cr), as well as for IMI-834 alloy (Ti—5.8Al—4Sn—3.5Zr—0.7Nb—0.5Mo—0.3Si). Mechanical testing of welded samplesfrom Ti-5553 alloy showed that joint ultimatestrength was equal to 94 % of BM strength(1042 MPa), yield point was 96 % (1005 MPa),relative elongation was 36 % of BM level. Low-ering of microhardness in the joint zone by morethan 17% compared to BM level at up to ±1.6 mmdistance from weld axis is found. Thermocoupleswere used to record the maximum temperatureof friction surface at LFW of IMI-834 alloy in

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the temperature range of 750—800 °C. It is shownthat for IMI-834 alloy the nature of dependenceof microhardness distribution on welding pa-rameters, similar to that for Ti-64 alloy is pre-served.

Work [19] is a study of LFW of titanium alloyTi—6Al—2Sn—4Zr—6Mo (Ti-6246). Depending onheat treatment, this alloy can be regarded as β-alloy in one case, and as α +β alloy in anothercase. Investigations were performed on samplesof 60 × 36 × 15 mm size. Welding of β-alloy andα + β-alloy Ti-6246 was performed. Postweldheat treatment included heating up to 620 °C andsoaking for 4 h. At certain welding parameterswelded joint ultimate strength (1078 MPa) onthe level of 96 % of BM values of the less strongα + β Ti-6246 alloy, and yield point (1030 MPa)equal to 99 % of this value for α + β Ti-6246alloy, are reached.

In [20, 21] formation of dissimilar joints atLFW of titanium alloys VT6 and VT8-1 was stud-ied. Welding was performed on 35 × 26 × 13 mmsamples with specifying 2 mm upsetting. Me-chanical properties of the joints were as follows:990—1020 MPa ultimate strength, 924—939 MPayield point, 11.39—13.78 % relative elongationwith fracture running through VT6 BM [21]. Attesting of special samples with a recess along thejoint line [20] for static tension, fracture ranthrough the thermomechanical impact zone(TMIZ) of VT8-1 alloy (ultimate strength of1194 MPa). Here, the width of the joint TMIZwas not larger than 400 μm, its greater part beingon VT8-1 alloy side.

Starting from the year 2000, the number ofpublications devoted to investigation of LFW ofhigh-temperature nickel super alloys becamemuch greater. Studied range of variation of superalloy LFW parameters was as follows: 2—3 mmamplitude, 40 to 100 Hz frequency, and 50—450 MPa pressure.

In [22] the process of LFW of Waspalloynickel alloy on 18 × 13 × 11 mm samples wasstudied. It is established that producing soundjoints is possible at a certain, high enough valueof all LFW technological parameters (amplitude,frequency, pressure). It is shown that at 0.9 mmdistance from the friction surface (dynamic re-crystallization zone – DRZ), metal temperaturewas not higher than 1126 °C, that led to disso-lution of strengthening γ′-phase and microhard-ness lowering. At 3.3 mm distance on both sidesof the weld microhardness lowering by 40 % com-pared to BM values is observed. Producing soundjoints (without pores or oxide inclusions) isachieved at more than 1.2 mm upsetting.

Work [23] is a microstructural study of jointsof IN-738 nickel alloy. Samples of 12.8 × 11.1 ×× 17.7 mm size were used for LFW. Completedissolution of strengthening γ′-phase to the depthof 600 μm on both sides of the weld is observed inthe joint zone. It is established that peak tempera-ture during welding was higher than 1230 °C.

In [24] optimization of LFW parameters forIN-718 nickel alloy was performed. Under certainconditions (for this alloy: 2 mm amplitude, 80 Hzfrequency, 70 MPa pressure), a temperature ofabout 1200 °C is reached on the friction surface.TMIZ is characterized by microhardness loweringby 25 % compared to BM. Microscopic studiesshowed presence of dispersed oxides across theentire weld section, dimensions and amount ofwhich depend on welding mode parameters.

Work [25] gives the results of studying LFWof CMSX-486 superalloy. Unlike titanium al-loys, where flash has the form of «petals» com-mon for both parts, for nickel alloys flash distri-bution into two individual «petals» is observedthat is characteristic for joints at rotational FW.At tensile testing samples failed through BM. Itshould be also noted that conducted in [23, 25]microstructural analysis of welded samples of IN-738 nickel superalloy and CMSX-486 single-crys-tal superalloy showed that melt areas can formin the alloys during LFW process. It is estab-lished that application of compressive load at theforging stage leads to rapid solidification of theformed metastable liquid phase.

Authors of [26] studied the applicability ofLFW for EP742 nickel alloy on samples of 35 ×× 26 × 13 mm size. Ultimate strength of weldedjoints was equal to 96—112 % of the value requiredby TU specification, and relative elongation was98—132 %. Width of DRZ and TMIZ was equalto 0.6 and 0.8 mm, respectively. Microhardnessdrop in the joint zone is observed that is due todegradation of strengthening γ′-phase.

LFW can be used to produce sound CM joints.In [27] joints of CM based on aluminium alloyAA2124 (Cu — 3.8, Mn — 0.5, Mg — 1.4, Zn << 0.25) reinforced by 25 % of silicon carbideparticles were studied. Sample size was 80 ×× 36 × 15 mm. Welding mode parameters wereas follows: 50 Hz vibration frequency, 2 mm am-plitude, 185 MPa heating and forging force, and2 mm upsetting. Mechanical testing of samplesshowed that ultimate strength of welded joint isequal to 82 % of BM level, proof strength is 78 %and relative elongation is 60 % of BM value.Microhardness lowering in the welding zone by10 % compared to BM of 2124Al/25 vol.% SiCcomposite is observed.

28 12/2014

Thus, as shown by analysis of publications,LFW application for nickel and titanium alloysis cost-effective, as here the cost of developmentof specialized equipment for aircraft engine con-struction products is justified. Weldability of ti-tanium alloys was studied in similar and dissimi-lar combinations, whereas LFW capabilities forsuperalloy welding are limited to investigationof joints in similar combinations.

Welded joint structure. On the whole, metalstructure at LFW is similar to that at rotationalFW [9]. Typical welded joint structure at LFWis given in Figure 4 for the case of AISI 316Lsteel [16].

At LFW metal in the joint zone is heated upto temperatures not exceeding the melting tem-perature. However, owing to heat conductivityand external pressure blank metal changes itsproperties and structure to a certain depth fromthe friction surface. As a rule, four zones differingby texture and microstructure, are distinguishedin the welded joint [9, 16, 28] as follows:

1. DRZ, or fine-grain zone is the welded jointcentral part. Here, the metal undergoes phasetransformations – an equiaxed fine-crystallinedynamically recrystallized structure forms;

2. TMIZ is characterized by metallographictexture. Grains and non-metallic inclusionstringers are elongated in the direction of defor-mation. In some materials phase transformationsin this zone can also take place;

3. HAZ is the region between BM and TMIZ,where phase or structural transformations canproceed, unrelated to deformation process;

4. BM zone is the region where heating duringwelding did not have any noticeable influenceon microstructure and mechanical properties.

Microstructural studies of samples of titaniumalloys Ti—6Al—4V in [18] and Ti—6Al—2Sn—4Zr—6Mo in [19] welded by LFW showed that DRZand TMIZ are directly proportional to weldingtime and inversely proportional to axial pressure,vibration frequency and amplitude. It is estab-lished that axial pressure is a determinant factorinfluencing DRZ and TMIZ dimensions.

For joints of similar materials structural changesare symmetrical relative to weld axis, unlike thosefor dissimilar material joints [20, 28].

LFW process modelling. In a number of stud-ies an attempt has been made to apply calculationmethods to describe energy characteristics ofLFW process and determine the temperaturemode of welding of various materials. In [29] itis proposed to determine the temperature in thebutt at LFW at the moment before upsettingapplication from the following equation:

T = T0 + 1

cρFn~√⎯⎯⎯⎯4πa ∫ 0

tq(τ)

√⎯⎯th — τ ×

× exp ⎛⎜⎝

⎜⎜—

z2

4a(t — τ) — b(th — τ)

⎞⎟⎠

⎟⎟ dτ,

(2)

where q(τ) is the law of variation of evolvedpower in time; th is the heating time; Fn~ is therod cross-sectional area; a is the thermal diffusiv-ity factor; b is the thermal efficiency factor; cρis the volumetric heat capacity; T0 is the initialtemperature level.

Instant thermal power q [7, 10] evolved atfriction can be determined from the followingformula:

q = Ffrv, (3)

where Ffr is the sliding friction force equal toFfr = Fkfr; (4)

where F is the welding force; kfr is the frictioncoefficient; v is the sliding velocity (velocity ofrelative displacement of parts).

Sliding velocity [24, 27] depends on ampli-tude α and frequency f of reciprocal motion. Foramplitude variable by sinusoidal law, averagesliding velocity vsl is equal to

vsl = 4αf. (5)

For sinusoidal variation of amplitude, slidingvelocity varies continuously from zero at peakamplitude up to maximum value at crossing apoint, in which the blanks being welded are co-axial. Pressure in the contact zone also varies –at peak amplitude it has the maximum value,

Figure 4. Microstructure of welded joint of AISI 316L steel

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and at zero crossing it has minimum value (cor-responds to welding force).

To determine the total energy [10], requiredfor welded joint formation, the expression forinstant thermal power is integrated over weldingprocess time t:

Ex = ∫ 0

t

qdt = ∫ 0

t

Ffrvdt.(6)

Friction force depends on a whole number offactors [1, 2]: velocity of relative motion of frictionsurfaces; material nature and presence of surfacefilms; temperature mode; normal pressure value;friction assembly rigidity and elasticity; stationarycontact duration; load application speed nature ofbody contact; surface size; relative overlap coeffi-cient; surface quality and roughness.

It is established [1] that at friction of steelsliding over steel, friction coefficient can vary inbroad ranges (0.1—1.0, depending on weldingconditions). Work [30] gives the data on frictioncoefficient variation during LFW for titaniumalloys in the range from 0.25 up to 0.55.

High intensity of physico-chemical processesproceeding in the joint zone at LFW makes mathe-matical modelling of this welding process a com-plex task, which is solved with application of com-puters and special software based on finite elementmethod. Simplified models are used to reduce thetime and computational power, for instance, cal-culation of temperature fields without allowing formaterial deformation, application of simplified 2D-models instead of 3D-modelling, consideration not

of the process as a whole, but of its individualstages; and prediction of the structure of indi-vidual welded joint zones [29—34].

Figure 5 gives the result of 2D-modelling ofLFW process. Authors of [29—34] performed cal-culation of temperature fields at LFW of titaniumalloys. It is established that for the studied pa-rameters of welding VT6 alloy about 30 % ofthermal energy, generated in the butt during theprocess, is removed into the flash together withplasticized metal. Calculations show that duringwelding temperature on friction surface reaches1300 K, it, however, can go up even to 1500 K,depending on LFW parameters. Here, tempera-ture increase up to 1000 °C lasts for less than 1 s[31]. In [33] temperature fields were calculatedfor welded joint of VT6 + VT8M-1 alloys; tem-perature field asymmetricity relative to the jointline is noted.

In [31, 32] flash formation was modeled inaddition to temperature fields. It is establishedthat flash topology depends on welding modeparameters. Producing sound joints requires com-plete renewal of the friction surfaces and removalof contaminations into flash.

Upsetting speed depends on pressure, vibra-tion amplitude and frequency. Their increase re-sults in higher deformation and upsetting speeds,reduction of plasticized metal zone thickness anddecrease of upsetting value required for surfacecleaning. Increase of friction speed leads tohigher maximum temperature in the butt, andwith increase of welding pressure it decreases.

Authors of [31] suggested a model for predic-tion of HAZ structure in welded joint of Ti-64alloy, as this zone is characterized by minimumstrength and hardness. Size of β-phase grains isthe main value of mechanical characteristics inthis alloy HAZ. Therefore, this model is basedon calculation of the size of β-phase grains afterLFW, depending on welding mode and grain sizebefore the process start.

In [30] modelling of LFW of links of round-link chains from 30CrNiMo8 steel was performed.Temperature variation during welding was con-sidered in two points removed for 3.5 and 4.5 mmfrom the butt. Data obtained by calculations dif-fer only slightly from experimental data. Differ-ence between calculated and experimental dataon sample upsetting was equal of 13 %.

LFW process modelling is limited to mostlytitanium alloys that is due to the urgency of thiswelding process application for manufacture ofcomponents of aircraft GTE. Joint formation atLFW of nickel superalloys, Ni—Al, Ti—Al basedintermetallics and CM in similar and dissimilarcombinations is not well enough highlighted in

Figure 5. Results of 2D-modelling of Ti-64 alloy for low(a), medium (b) and high (c) welding pressures [34]

30 12/2014

publications that necessitates performance of fur-ther research.

Conclusions

1. Requirements of leading enterprises in aircraftpropulsion engineering determined the urgencyof development of technologies and equipmentfor LFW of parts from high-strength and high-temperature alloys. LFW allows producingsound joints of titanium alloys, different steels,high-temperature nickel alloys and CM.

2. Structure of joints in LFW is similar tothat for other FW variants. Width of charac-teristic sections of the joint (zones of dynamicrecrystallization, thermomechanical and thermalimpact) depends on LFW mode parameters,namely axial force, welding time, vibration am-plitude and frequency.

3. Mathematical modelling of LFW process,based on finite element method, allows assess-ment of thermodeformational conditions of jointformation and prediction of structural changesin materials in the welding zone.

4. Lowering of the cost and improvement ofreliability of equipment for realization of LFWof high-temperature and high-strength materialsremains to be a problem, limiting the field ofLFW application to predominantly aerospace in-dustry enterprises.

5. Performance of further studies on LFW ofnickel superalloys, Ni—Al, Ti—Al based intermet-allics and CM in similar and dissimilar combina-tion is urgent.

1. Lebedev, V.K., Chernenko, I.A., Vill, V.I. et al.(1987) Friction welding: Refer. Book. Leningrad:Mashinostroenie.

2. Vill, V.I. (1962) Friction welding of metals. NewYork: American Welding Soc.

3. Vavilov, A.F., Voinov, V.P. (1964) Friction weld-ing. Moscow: Mashinostroenie.

4. Maurya, R., Kauzlarich, J. Bonding apparatus –friction welding by reciprocal motion. Pat. US3420428-A. Publ. 1969.

5. Kulis, E.I., Lokishin, R.F. (1982) Friction weldingof plastics. Svarochn. Proizvodstvo, 1, 8—9.

6. Nicholas, E.D., Hone, P. (1989) Developments infriction and MIAB welding. Welding Institute Bull.,R382/11/89.

7. Nicholas, E.D. (1991) Linear friction welding. Dues-seldorf: DVS Verl., DVS Berichte, 139, 18—24.

8. Nicholas, E.D. (1992) Friction surfacing and linearfriction welding. In: Proc. of Int. SAMPE and Met-als Proc. Conf. (Covina, CA, USA, 1992), 450—463.

9. Bhamji, I., Preuss, M., Threadgill, P.L. et al. (2010)Solid state joining of metals by linear friction weld-ing: A literature review. Mater. Sci. & Technol.,27(1), 2—12.

10. Usani, U. Ofem, Colegrove, P.A., Addison, A. et al.(2010) Energy and force analysis of linear frictionwelds in a medium carbon steel. Sci. and Technol. ofWelding & Joining, 15(6), 479—485.

11. Vairis, A., Frost, M. (1998) High frequency linear fric-tion welding of a titanium alloy. Wear, 217, 117—131.

12. Wanjara, P., Jahazi, M. (2005) Linear friction weld-ing of Ti—6Al—4V: Processing, microstructure andmechanical property interrelationships, metal. Mater.Transact. A, 36, 2149—2164.

13. Netwig, A. (1993) Entwicklung und Trends beimReibschweissen. Der Praktiker, 9, 546—555.

14. Koo, H.H., Baeslack, W.A. (1992) Friction weldingof rapidly solidified Al—Fe—V—Si alloy. Welding J.,5, 20—24.

15. Addison, A., Threadgill, P. (2010) Initial studies oflinear friction welding of C—Mn steel. Welding andCutting, 4, 364—370.

16. Bhamji, I., Preuss, M., Threadgill, P.L. et al. (2010)Linear friction welding of AISI 316L stainless steel.Mater. Sci. and Eng., 528, 680—690.

17. Kallee, S.W., Nicholas, E.D., Russell, M.J. (2003)Friction welding of aeroengine components. In: Proc.of 10th World Conf. on Titanium (Hamburg, Ger-many, 2003), 2859—2867.

18. Evolution of microstructure, microtexture and mechani-cal properties in linear friction welded titanium alloys.digitool.library.mcgill.ca/thesisfile 103485. pdf

19. Corzo, M., Torres, Y., Anglada, M. et al. (2007)Fracture behaviour of linear friction welds in tita-nium alloys. Annales de la Mecanica de Fractura, 1,75—80.

20. Diakonov, G.S., Izmajlova, N.F., Bychkov, V.M. etal. (2012) Examination of weld zone microstructurein linear friction welding of VT6 and VT8 titaniumalloys. Vestnik UGATU, 16(7), 48—52.

21. Medvedev, A.Yu., Bychkov, V.M., Selivanov, A.S. etal. (2012) Application of linear friction welding forjoining of VT6 and VT8-1 alloys. Ibid., 16(7), 63—67.

22. Chamanfar, A., Jahazi, M., Gholipour, J. et al.(2011) Mechanical property and microstructure oflinear friction welded WASPALLOY. Metallurg.and Mater. Transact. A, 42(March), 729—744.

23. Ola, O.T., Ojo, O.A., Wanjara, P. et al. (2012)Analysis of microstructural changes induced by linearfriction welding in a nickel-base superalloy. Ibid., 42(Dec.), 3761—3777.

24. Linear friction welding of IN-718: Process optimiza-tion and microstructure evolution http://www.sci-entific.net/AMR.15-17.357.

25. Ola, O.T., Ojo, O.A., Wanjara, P. et al. (2012) Astudy of linear friction weld microstructure in singlecrystal CMSX-486 superalloy. Metallurg. and Ma-ter. Transact. A, 43(March), 921—933.

26. Bychkov, V.M., Selivanov, A.S., Medvedev, A.Yu.et al. (2012) Study of weldability of heat-resistantnickel alloy EP742 by method of linear friction weld-ing. Vestik UGATU, 16(7), 112—116.

27. Rotundo, F., Ceschini, L., Morri, A. et al. (2010)Mechanical and microstructural characterization of2124Al/25 vol.% SiCp joints obtained by linear fric-tion welding (LFWptA). Composites, 41, 1028—1037.

28. Karavaeva, M.V., Kiselyova, S.K., Bychkov, V.M. etal. (2012) Influence of upset value on formation ofwelded joint in linear friction welding. Pisma o Ma-terialakh, 2, 40—44.

29. Medvedev, A.Yu., Pavlinich, S.P., Atroshchenko, V.V.et al. (2010) Modeling of temperature field in linearfriction welding. Vestik UGATU, 14(2), 75—79.

30. Wen-Ya Li, Tiejun Ma, Jinglong Li. (2010) Numeri-cal simulation of linear friction welding of titaniumalloy: Effects of processing parameters. Materialsand Design, 31, 1497—1507.

31. Grujicic, M., Arakere, G., Pandurangan, B. et al.(2012) Process modeling of Ti—6Al—4V linear frictionwelding (LFW). J. Mater. Eng. and Performance,21(10), 2011—2023.

32. The importance of materials data and modeling pa-rameters in an FE simulation of linear friction weld-ing. http://www.hindawi.com/journals

33. Kiselyeva, S.K., Yamileva, A.M., Karavaeva, M.V.et al. (2012) Computer modeling of linear frictionwelding based on the joint microstructure. J. Eng.Sci. and Technol. Rev., 5, 44—47.

34. Turner, R., Gebelin, J.-C., Ward, R.M. et al. (2011)Linear friction welding of Ti—6Al—4V modeling andvalidation. Acta Materialia, 59, 3792—3803.

35. Linear friction welding of high strength chains.http://www.raiser.de/download/innovationspreis/bewerber2013/Linear_Friction_Welding_of_High_Strength_Chains_Mucic-Fuchs-Enzinger.pdf

Received 09.10.2014

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FLASH-BUTT WELDING OF THIN-WALLED PROFILESOF HEAT-HARDENED ALUMINIUM ALLOYS

P.N. CHVERTKO, L.A. SEMYONOV and K.V. GUSHCHINE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

Thin-walled profiles of high-strength aluminium alloys have found application in longitudinal load-carryingelements of flying vehicles and other critical structures. In industry, the serial profiles, heat-treated formaximum strength, are used, that excludes the feasibility of their welding by fusion arc methods. Theproblem of producing the welded joints of the above-mentioned load-carrying elements in solid phase isurgent. The present work is aimed at the study of formation of thin-walled profile joints of dissimilar(AK6 + D16) and similar (V95) heat-hardened aluminium alloys in the flash-butt continuous welding.Peculiarities of the resistance flash-butt welding with formation of joints during upsetting with extrusionwere investigated and basic technology was developed, which is used for producing joints of aluminiumalloy parts. This technology allows increasing greatly the quality of welded joints of this group of alloys,and also widening the range of thicknesses of metal, being joined by the flash-butt welding. Given are themain parameters of the flash-butt continuous welding modes. Strength properties of welded joints are notless than 90 % of the base metal strength. 6 Ref., 8 Figures.

Keywo r d s : flash-butt welding, continuous flashing,upsetting, formation of joints, aluminium alloys

Alloys V95, D16 and AK6 are widely used inmanufacture of critical elements and structuresin aircraft and rocket construction. Thus, at thepresent time the riveted joints of alloys V95,

AK6 + D16 are used for manufacture of a longi-tudinal load-carrying structure of casings of aero-space engineering equipment, because these al-loys are referred to the hard-to-weld group. Theimportant drawback of the riveted joint is theincrease in mass of structure due to the appear-ance of auxiliary elements during riveting of ele-ments being abutted (Figure 1, a). The rivetingis a labor-consuming operation, connected withhard labor conditions. It is necessary to make acareful treatment of the hole surface for rivetingto provide the reliable joining of the product.During long-term service the riveted joint is sub-jected to loosing, having an influence on theproduct service life.

Application of welding instead of riveting(Figure 1, b) is one of the effective methods ofsolution of the problem of improving thestrength, quality of the joint and increasing thetactical-technical characteristics of flying vehi-cles, in particular, the decrease in structure massand, accordingly, the increase in payload of theflying vehicles [1].

One of the main load-carrying elements of theflying vehicle structure is the stringer-fittingjoints. Stringers are manufactured in majority ofcases from T-section of high-strength heat-hard-ened alloys D16, V95, etc. Fittings are manufac-tured by milling from different semi-products(for example, forgings of alloy AK6 or plates ofthe same alloy as the stringer panels).

In this connection the assessment of weldabi-lity of thin-walled profiles of different sections

Figure 1. Joining of elements of fitting-stringer type oflongitudinal load-carrying structure of flying vehicles byriveting (a) and flash-butt welding (b)

© P.N. CHVERTKO, L.A. SEMYONOV and K.V. GUSHCHIN, 2014

32 12/2014

of alloy V95-T1 and plates of alloy D16-T + AK6-T1 was made during their welding.

Semi-products of alloy D16 and V95 havefound a wide application in industry. These alloysare hardened by the heat treatment and, as aresult, acquire the high mechanical propertiesand retain a sufficient technological ductility.

Among the high-strength alloys on aluminiumbase alloy V95, which is referred to Al—Zn—Mg—Cu (5.0—7.0 % Zn, 1.8—2.8 % Mg, 1.4—2.0 % Cu), found the largest application inrocket-space and aircraft structures. Zinc, mag-nesium and copper form solid solutions with alu-minium and between themselves and differentmetallic compounds – phases M (MgZn2), S(Al2CuMg), T (Al2Mg3Zn3), playing a large rolein alloy hardening at its heat treatment.

D16 is the alloy of, at least, six components:aluminium, copper, magnesium, manganese, ironand silicon, though the main alloying elementsare copper and magnesium (3.8—4.9 % Cu, 1.2—1.8 % Mg), therefore it refers to the alloys ofAl—Cu—Mg system.

Alloy AK6 of Mg—Si—Cu alloying system ismainly used in the form of forgings, producedmainly from the pressed rods. Main alloying ele-ments are magnesium, silicon and copper (0.6—1.0 % Mg, 0.9—1.0 % Si, 2.0 % Cu). The alloyis widely used in industry (construction, trans-port machine building, aviation) for manufactureof stamped and forged parts of intricate shape,as well as for loaded parts of the type of frames,fittings, etc. Microstructure of alloy AK6 consistsafter heat treatment of grains of aluminium solidsolution and inclusions of metallic compoundsCuAl2 and Mg2Si. Alloy AK6 is less sensitive toheating than alloys D16 and V95 [2, 3].

Thermomechanically strengthened alloys arevery sensitive to heating. Degree of weakeningdepends on heating temperature and time of heat-ing duration. During welding the highest me-chanical properties can be obtained in that casewhen the duration of heating up to the tempera-tures above critical ones does not exceed the defi-nite limits [4]. It is difficult to provide suchtemperature cycle in welding of aluminium alloysdue to their heat conductivity. The intensivehigh-concentrated heat input to the heating zoneis required.

One of the most challenging methods of pro-ducing the quality welded joints with high me-chanical properties is the flash-butt welding(FBW) with a continuous flashing.

The present work is aimed at the study ofpeculiarities of formation of joints of thin-walledprofiles of dissimilar (D16 + AK6) and similar

(V95) heat-hardened aluminium alloys at thecontinuous FBW with extrusion.

The welding process is performed automat-ically, providing high stable quality of the joint.Design of welding equipment and technologicalrigs provide combining of assembly-weldingworks in a single cycle and high accuracy of thewelded joint geometric sizes [5].

The obligatory condition of producing thequality welded joints of aluminium alloys is theformation of joint with metal extrusion duringupsetting into gap between the forming devices.In this case, the degree of deformation is in-creased with knives approaching [6]. Scheme ofFBW process with the joint formation is givenin Figure 2. The forming devices are fulfillingtwo important functions: during upsetting theyform the joint under conditions of volume-plasticdeformation, and also fulfill the function ofknives for the flash cutting. As a result thewelded joint is formed, which does not need thenext mechanical cleaning from the flash (Fi-gure 3).

Welding of alloys of different systems of al-loying (D16-T and AK6-T1) was carried out onplates of 2.5—5.0 thickness and 25—35 mm width.

To make prototyping of the welded joint ofelements of the fitting-stringer type, the inves-tigations were carried out on T-section of alloy

Figure 2. Scheme of FBW with formation of joint: 1 –parts; 2 – forming devices; 3 – current connector; 4 –extruded metal; lw – tolerance for welding

Figure 3. Welded joint of T-section profile of aluminiumalloy V95-T1

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V95-T1 with thickness of flanges δ = 2.5 and4.0 mm. Welding fixture in the form of jigs wasdesigned for the experiments on welding.

Welding of specimens was carried out in theuniversal laboratory machine for FBW with con-tinuous flashing, which was equipped by apneumo-hydraulic drive of upsetting at the forceFups = 130 kN and welding transformer of150 kV⋅A capacity.

Differences in physical properties of alloys(heat conductivity of alloy AK6 is by 30 % higherthan that of alloy D16 [2, 3]) have a great in-fluence on their heating during flashing, weldformation, character of deformation and struc-tural transformations in the HAZ.

Selection and subsequent correction of weld-ing modes were carried out experimentally withaccount for above-mentioned peculiarities of al-loys. Express analysis of the welded joint quality,consisting in bending of notched specimens acrossweld up to fracture, was carried out in the processof run-out of the welding modes. The joint qual-ity was evaluated by the presence or absence ofdefects in visual inspection of the specimen, frac-tured across weld (oxide films, etc.). Attemptswere made to minimize the time of welding inwelding modes to reduce the metal weakening

during heating. Using the described procedure,the optimum modes, which provide the absenceof defects in fusion line, were determined.

Main parameters of welding: secondary volt-age U2o.c = 4 V, rate of flashing is changed ex-ponentially from 2 up to 25 mm/s, upsettingrate is not more than 5 s, time of welding is notmore than 5 s.

The distribution of temperature in near-con-tact zone of edges heated by flashing directlybefore upsetting is given in Figure 4.

To determine the redistribution of alloyingelements between alloys in weld, the X-ray spec-tral microanalysis of AK6-T1 + D16-T weldedjoints was made (Figure 5). Transition zone ofredistribution of alloying elements in weld is ap-proximately 300 μm. Beyond the transition zonethe content of alloying elements in weld corre-sponds to the base metal composition.

In the process of upsetting the texture ofwelded joint with a typical turning of base metalfibers at 90° is formed as a result of extrusion ofmetal between the forming devices. Figure 6 gives

Figure 4. Distribution of temperature for depth l heated byflashing of near-contact zone 4.0 mm thick plate of alloyD16-T before upsetting

Figure 5. Distribution of elements in welded joint (a) andHAZ (b)

Figure 6. Macrostructure of welded joint butt: a – AK6 + D16; b – V95

34 12/2014

the microstructures of welded joints. The micro-structure of the joint zone is characterized by thedeformed elongated grains of solid solution at ahigh density of intermetallic phases in the formof elongated chains (Figure 7). With weld ap-proaching the intermetallic inclusions are re-fined. In the structure of welded joint profile ofalloy V95-T1 the weld 6.0 μm thick is formed,which has a typical fine-grained structure of solidsolution and refined intermetallic phases of 1.0—2.0 μm.

According to the results of carried out inves-tigations of the distribution of metal hardness inwelded joints (Figure 8) the total HAZ value is15—20 mm, maximum weakening of weld doesnot exceed 6 %.

Results of mechanical tests of the base metaland welded specimens of AK6-T1 + D16-T alloysshowed the strength of the welded joints at thelevel of base metal of the less strong alloy AK6-T1(σt

w = 386 MPa, σtBM AK6 = 387 MPa, σt

BM D16 == 455 MPa). Specimens were fractured in HAZon the side of alloy AK6-T1.

The strength of welded joints of V95-T1 alloyprofile is at the level higher than 90 % of thebase metal strength ((σt

w = 521—542 MPa,σt

BM B95−T1 = 580 MPa)

From the obtained results of investigationsthe positive conclusion can be made about thechallenging application of FBW technology withcontinuous flashing for joining the elements ofthin-walled profiles of high-strength aluminiumalloys in critical structures. This technologymakes it possible to weld elements of structuresin heat-hardened state with a loss in strength ofnot more than 10 %.

1. Nikolaev, G.A., Fridlyander, I.N., Arbuzov, Yu.P.(1990) Aluminium alloys to be welded. Moscow:Metallurgiya.

2. (1973) Aluminium alloys. Application of aluminiumalloys: Manual. Ed. by A.T. Tumanov, I.N., Fridlyan-der. Moscow: Metallurgiya.

3. (1984) Commercial aluminium alloys: Refer. Book.Ed. by F.I. Kvasov, I.N. Fridlyander. Moscow: Me-tallurgiya.

4. Kuchuk-Yatsenko, S.I., Chvertko, P.N., Semyonov,L.A. et al. (2010) Peculiarities of flash butt weldingof high-strength aluminium alloy 2219. The PatonWelding J., 3, 7—9.

5. Kuchuk-Yatsenko, S.I. (1992) Flash-butt welding.Kiev: Naukova Dumka.

6. Kuchuk-Yatsenko, S.I., Chvertko, P.N., Semyonov,L.A. et al. (2013) Flash butt welding of products ofhigh-strength alloys based on aluminium. The PatonWelding J., 7, 2—6.

Received 19.05.2014

Figure 7. Microstructure of weld alloy AK6-T1 + D16-T (a) and V95-T1 (b) joints

Figure 8. Distribution of hardness in welded joint of plates from alloys AK6-T1 + D16-T and profile of alloy V95-T1 (b)

12/2014 35

APPLICATION OF COMPLEX-ALLOYED POWDERSPRODUCED BY THERMOCENTRIFUGAL SPUTTERING

IN FLUX-CORED WIRES

A.P. ZHUDRA, S.Yu. KRIVCHIKOV and V.I. DZYKOVICHE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

This paper presents the results of evaluation of the influence of complex-alloyed powders (CAP) onwelding-technological properties of flux-cored wires and deposited metal performance. Possibility of ensuringa high degree of flux-cored wire alloying at reduction of its diameter is considered. Tribotechnical charac-teristics of metal deposited with flux-cored wires are determined at CAP application as a filler. The paperpresents the results of metallographic investigations of deposited metal. It is established that CAP applicationas flux-cored wire core is favourable for the mode of electrode metal melting and transfer leading toimprovement of their welding-technological characteristics. Wear resistance studies of deposited metalshowed an improvement of wear resistance of deposited metal operating under the conditions of abrasivewear and metal-over-metal friction. 4 Ref.. 5 Tables. 3 Figures.

Keywo r d s : flux-cored wire, thermocentrifugal sput-tering of ingots, complex-alloyed powders, welding-tech-nological characteristics of flux-cored wire, tribotechni-cal characteristics of deposited metal

Flux-cored wires produced to GOST 26101—84have a core, in which the alloying part consistsof crushed ferroalloys and metal powders. En-suring a high degree of alloying, usually, neces-sitates application of high (40—50 %) values ofthe filling coefficient. This, in its turn, leads toincrease of flux-cored wire diameter up to 3.2 mmand more. Stable process of surfacing with suchflux-cored wire requires considerable current(400—500 A) and arc voltage (28—32 V), that isnot rational or inadmissible for some types ofparts being reconditioned, which are sensitive tohigh temperature impact, prone to thermal de-formation, etc. (in such cases application of large-diameter flux-cored wires leads to considerablepower consumption). Thus, solution of the prob-lem of flux-cored wire diameter reduction at pres-ervation of its required alloying level, is of con-siderable practical interest.

PWI developed the technology of producingcomplex-alloyed powders (CAP), which consistsin melting out an ingot of cylindrical form of therequired composition from a mixture of ferroal-loys and other alloying materials, and its sub-sequent plasma-arc centrifugal sputtering in themodes, ensuring production of spherical powderparticles of the required granulometric composi-tion. The composition of each particle shouldcorrespond to calculated composition of surfacing

electrode materials, namely flux-cored wires,strips, filler rods, stick electrodes , etc.

In addition, spherical shape of CAP particles(Figure 1, b), unlike irregular-shaped particles(Figure 1, a), will ensure a more dense filling ofthe core and will enable reducing flux-cored wirediameter.

On the other hand, replacement of mechanicalmixture of core materials of nonuniform struc-ture, properties and chemical composition byCAP, will result in changes of core electric andthermophysical properties. This, in its turn, willinfluence the welding-technological properties offlux-cored wire and, possibly, deposited metalcharacteristics.

The objective of this work is experimentalinvestigation of CAP influence on welding-tech-nological characteristics of flux-cored wires andwear resistance of deposited metal. Earlier [1]PWI performed investigations on application ofgranulated alloy of the respective composition asthe core of PP-Np-25Kh5FMS flux-cored wire.This study focused on improvement of depositedmetal high-temperature resistance.

As objects of study, we selected two flux-cored wires of another type: PP-AN170 (PP-Np-80Kh20R3T), applied for hardfacing of parts op-erating under the conditions of intensive abrasiveand hydroabrasive wear (teeth of excavator la-dles, dredge noses, etc), and PP-AN160 (PP-Np-200KhGR) used for hardfacing of parts operatingunder the conditions of lubricated friction of met-al over metal (cast iron crankshafts of car en-gines).© A.P. ZHUDRA, S.Yu. KRIVCHIKOV and V.I. DZYKOVICH, 2014

36 12/2014

Two pairs of flux-cored wires were manufac-tured to perform comparative testing: first –PP-AN170 and PP-AN170SF, and the second –PP-AN160 and PP-AN160SF. In each pair thefirst flux-cored wire is manufactured in keepingwith the requirements of GOST 26101—84, andthe second – with application of CAP as thecore.

Composition of CAP, produced by thermocen-trifugal sputtering in OB-2327 unit [2], is givenin Table 1.

Modes of test ingot sputtering Welding current, A .......................................... 350—500Arc voltage, V ..................................................... 28—34Rotation speed, rpm ...................................... 3500—5000Ingot feed rate, m/min .................................... 1.50—1.5

To assess CAP influence on welding-techno-logical characteristics of test flux-cored wires,deposition of isolated beads was performed in thefollowing mode: Id = 280—300 A, Ua = 26—28 V,vd = 15 m/h for flux-cored wires PP-AN170 andPP-AN170SF, and Id = 170—180 A, Ua = 20—26 V,vd = 7.7 m/h for flux-cored wires PP-AN160 andPP-AN160SF. Kind of current is direct, polarityis reverse.

Conducted testing (Table 2) revealed that CAapplication as flux-cored wire core improves arc-ing stability (determined by the value of arc volt-age variation Kv: the smaller Kv value, the morestable the arcing process), reduces electrode met-al losses for sputtering ψsp, and promotes im-provement of the quality of deposited metal for-mation b/h (assessed by the value of the ratioof width b of deposited bead to its height h: thelarger b/h value, the higher the formation qual-ity). Spherical shape of CAP particles allowedreducing flux-cored diameter dfl.w and, thus, low-

ering minimum possible values of arc current andvoltage, at which the surfacing process runs in astable manner. CAP positive influence on Kv,b/h and ψsp is due to favourable nature of theirmelting and electrode metal transfer: more uni-form melting of the core and sheath, reductionof the size of transferring electrode drops, quan-tity and duration of short-circuiting of the arcgap. The most significant changes of the aboveparameters are found at CAP application as thecore of PP-AN170 flux-cored wire.

Figure 1. Appearance of fragmented particles of ferroalloys(a) and spherical granules produced by the method of ther-mocentrifugal sputtering (b)

Table 1. Composition of CAP for flux-cored wire production

Flux-cored wire typeElement content, wt.%

C Cr Mn Si Ti B Al Fe

PP-AN170 3 50 4 4 1.5 8 — —

PP-AN160 19 5 9 13 5 1 2 Res.

Table 2. Welding-technological characteristics of test flux-cored wires

Flux-cored wire type

Welding-technological parameter values

ψsp, % b/h Kv, %Surfacing modes and flux-cored wire diameters

Id, A Ua, V dfl.w, mm

PP-AN170 16.8 2.0 25.6 280—300 26—28 3.2

PP-AN170SF 9.2 4.3 18.5 220—240 22—24 2.6

PP-AN160 7.4 4.5 9.4 170—180 21—22 1.8

PP-AN160SF 6.2 5.0 7.2 140—150 18—20 1.6

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Abrasive wear-resistance testing of metal de-posited with PP-AN170 and PP-AN170SF flux-cored wires was conducted in NKM friction ma-chine by «stationary ring» schematic at frictionagainst non-fixed abrasive [3]. Test samples of16 × 16 × 6 mm size were cut out of the third orfourth layer of multilayer deposit, produced inthe above mode. Deposited metal composition isgiven in Table 3. Relative wear resistance wasassessed as the ratio of the loss of reference-sam-ple mass (metal deposited with flux-cored wirePP-AN170) to loss of mass of the sample, depos-ited with PP-AN170SF flux-cored wire. Derivedresults are given in Table 4.

Test results indicate that CAP application asthe core of PP-AN170 flux-cored wire leads toincrease of deposited metal abrasive wear resis-tance by 22—23 %. Comparative metallographicanalysis of deposited metal structures (Figure 2)showed that CAP does not have any influence

on composition, but changes the degree of dis-persity and hardness of structural components.

So, metal deposited with PP-AN170SF flux-cored wire (Figure 2, b) is characterized by for-mation of carboborides and eutectic with higherdegree of dispersity than for metal deposited withPP-AN170 flux-cored wire (Figure 2, a). Bothin the first and in the second case hardness ofcarboborides and eutectic is practically the same,and is equal to HV1-14000—17000 and HV1-9200—14500, respectively. At the same time,hardness of the matrix of metal deposited withPP-AN170CF flux-cored wire is higher and isequal to HV1-7200—8600, whereas for metal de-posited with AN170 flux-cored wire it is HV1-6500—7400. In all probability, higher hardnessof deposited metal matrix is also responsible forits higher abrasive wear resistance.

Determination of tribotechnical parameters ofmetals deposited with PP-AN160 and PP-AN160SF flux-cored wires was conducted in fric-tion machine SMT-1 by roller—block method toGOST 23224—86. Blocks were segments of stand-ard car inserts AO 20-1. Blanks of VCh50-2 castiron were used for roller production. Their cy-lindrical surface was surfaced with test flux-cored wires in the above modes, and then groundto nominal dimensions: 12 mm thickness and50 mm diameter. Composition of tested depos-ited metals is given in Table 3.

Testing of friction-sliding pairs with lubrica-tion was performed in two stages: in running-inmode and at working load. Value of optimumworking load Pop was selected on the basis ofpreliminary experiments, which showed that itsincrease can lead to «seizure» on deposited met-al—insert mating surface. Total wear rate IΣ onthe whole was determined as a sum of intensitiesof wearing of insert Ib and roller Ir. Procedureof determination of these values is given in GOST23.224—86. In addition, during testing the changeof the coefficient of mating surface friction ffr

Table 3. Chemical composition of deposited metal, wt.%

Flux-cored wiretype

C Cr Si Mn Ti Al B

PP-AN160 2.2 0.3 1.2 0.8 0.3 0.2 0.08

PP-AN160SF 2.0 0.5 0.8 0.6 0.3 0.1 0.1

PP-AN170 1.1 18.5 0.7 0.8 0.3 — 2.6

PP-AN170SF 0.9 19.1 0.5 0.8 0.4 — 2.9

Figure 2. Structure (×500) of metal deposited with flux-cored wires PP-AN170 (a) and PP-AN170SF (b)

Table 4. Results of abrasive wear resistance testing of metal de-posited with PP-AN170 and PP-AN170SF flux-cored wires

Samplegrade

Sample mass, gWear, g

Relativewear

resistanceBefore testing After testing

170-1 12.0436 11.9017 0.1419 1

170SF-1 12.0003 11.8888 0.1115 1.27

170-2 12.4266 12.2736 0.1530 1

170SF-2 11.6245 11.4991 0.1254 1.22

170-3 12.1111 11.9717 0.1394 1

170SF-3 12.2234 12.1176 0.1058 1.32

38 12/2014

and temperature of frictional heating of oil Tfrwas recorded.

Obtained data (Table 5) show that wear re-sistance of deposited metal of mating surfaces 1and 2 is approximately the same. At the sametime IΣ of mating surface 2 is 1.5 times lowerthan that of mating surface 1, that is due tosmaller I,b value of mating surface 2. Proceedingfrom that it can be stated that for the two com-pared mating surfaces application of PP-AN160SF flux-cored wire provides depositedmetal with higher antifriction properties.

Metallographic examination of deposited met-al of mating surfaces 1 and 2 was conducted todetect the causes affecting the derived results. Itshould be noted that detailed analysis of struc-tural-phase state of the multicomponent alloy,which solidified under non-equilibrium condi-tions, is a rather labour-consuming task. How-ever, it can be stated with sufficient degree ofvalidity that the deposited metal structure (atapplication of both PP-AN160 and PP-AN160SFflux-cored wire) consists of two main phases:austenite decomposition products (pearlite + re-sidual austenite) and carbide-cementite phase(Figure 3). The latter has the form of ramifiedreinforcing net with regions of different thicknessin the section plane.

In the micrograph this phase is of white colour,and pearlite-austenite phase is black that allowedassessment of their quantitative ratio. As a result,it was established that deposited metal of matingsurface 1 contains 40—46 % of «white» (carbide-cementite) phase of hardness HV-0.5 8000—8200.Hardness of solid solution grains is HV0.5-5000—5200. In the deposited metal of mating surface 2the fraction of carbide-cementite phase is 28—33 %, and its hardness remains to be the same asin the deposited metal of mating surface 1. Atthe same time, pearlite hardness in deposited met-al of mating surface 2 (HV-0.5-5800—6200) issignificantly higher than that of deposited metalof mating surface 1.

Thus, it can be stated that reduction of thefraction of hard carbide phase in the depositedmetal, which is also abrasive relative to the in-sert, promotes lowering of wear rate of the insert

and mating surface as a whole. At the same time,reduction of the quantity of wear-resistant struc-tural component is compensated by increase ofsolid solution hardness and quantity. As a result,wear rates of deposited metal of mating surfaces1 and 2 remain approximately the same. Metal-lographic analysis could not reveal the mecha-nism of CAP influence on quantitative ratio ofphases in the deposited metal and their hardness.It can be assumed, however, that this is relatedto influence of boron as an active modifying andcarbide-forming element.

In [4] it is shown that the form of boron-con-taining material (B4C, B2O3, BN, etc.) in the

Table 5. Tribotechnical characteristics of friction-sliding pairs

Mating surfacenumber

Type of mating surface andflux-cored wire

Performance characteristics Wear rate, mm/m⋅10—4

ffr⋅10—2

Pop, MPa Tfr, °C Ir Ib IΣ

1 Deposited metal—insert.PP-AN160 flux-cored wire

11 72 0.24 0.26 0.5 45

2 Deposited metal—insert.PP-AN160SF flux-cored wire

14 50 0.22 0.12 0.34 20

Figure 3. Structure (×320) of metal deposited with flux-cored wire PP-AN160 (a) and PP-AN160SF (b)

12/2014 39

composition of flux-cored wire core has an essen-tial influence on their quantitative ratio and mor-phology of deposited metal phase components.Even though boron was added to the compositionof PP-AN160 and PP-AN160SF flux-cored wirecores in the same amounts, but different boron-containing materials were used: in the firstcase – iron-chromium-boron master alloy, andin the second case – CAP particles. It is alsopossible that different content of carbide-cemen-tite phase in the studied samples of depositedmetal is the result of varying carbide-formingactivity of boron, depending on its compositionalstate.

Conclusions

1. Application of complex-alloyed powders ascore of complex-alloyed flux-cored wires allowsan essential reduction of its diameter withoutchanging the alloying level, reduction of powerconsumption in surfacing processes and improve-ment of deposited metal performance. Replace-ment of core of PP-AN170 flux-cored wire byCAP core allowed improvement of its welding-technological characteristics, reducing its diame-ter from 3.2 to 2.6 mm and increasing abrasivewear resistance of deposited metal by 1.2—1.3times.

2. Replacement of mechanical mixture ofcharge materials of non-uniform composition influx-cored wire core by CAP influences the quan-titative composition and hardness of depositedmetal structural components. CAP applicationin PP-AN160 flux-cored wire reduces the fractionof abrasive carbide phase and increases the hard-ness of deposited metal pearlite base that in-creases by 1.3 times the wear resistance of matingsurface exposed to metal-over-metal friction inservice.

1. Kondratiev, I.A. (1980) Flux-cored wire filled bygranular alloy. In: Theoretical and technologicalprinciples of surfacing. Surfacing of metallurgicaland power engineering parts. Kiev: PWI.

2. Dzykovich, V.I. (2010) Study and development ofmaterials for wear-resistant surfacing based onspheroidized granules of tungsten carbides: Syn. ofThesis for Cand. of Techn. Sci. Degree. Kiev.

3. Yuzvenko, Yu.A., Gavrish, V.A., Marienko, V.Yu.(1979) Laboratory units for evaluation of wear resis-tance of deposited metal. In: Theoretical and techno-logical principles of surfacing. Properties and testsof deposited metal. Kiev: PWI.

4. Zhudra, A.P., Krivchikov, S.Yu., Petrov, V.V.(2004) Selection of boron-containing charge materialsfor the core of flux-cored wire. The Paton WeldingJ., 4, 51—52.

Received 22.09.2014

40 12/2014

MODERNIZATION OF CONTROL SYSTEMOF A1756 MACHINE FOR PLASMA-POWDER SURFACING

E.F. PEREPLYOTCHIKOV, I.A. RYABTSEV, Yu.N. LANKIN, V.F. SEMIKIN and P.P. OSECHKOVE.O. Paton Electric Welding Institute, NASU

11 Bozhenko Str., 03680, Kiev, Ukraine. E-mail: office@paton.kiev.ua

Given are the results of modernization of control system of A1756 machine for plasma-powder surfacing.New control system allows for complete automating of surfacing process as well as optimizing consumptionof filler powder and minimizing its loss. Control of surfacing process is carried out with the help ofprogrammer, designed using PIC16F886 (Microchip) microcontroller and alphanumeric OLED indicatorWEH1601A (Winstar). Microprocessor block of the programmer allows for setting the parameters of timecharacteristics for surfacing current and powder consumption and their processing in process of programcontrol. Scheme of programmer connection to A1756 machine with capability of operation switching fromprogrammer as well as in simple design mode was developed and built. The programmer allows for carryingout setting of modes, providing good bead formation during automatic plasma-powder surfacing, qualityfilling of crater and minimum loss of filler powder. 2 Ref., 7 Figures.

Keywo r d s : plasma-powder surfacing, surfacing ma-chine, control system, powder consumption, surfacingmode, programmer

Characteristics of transfer and melting of fillermaterial in plasma-powder surfacing are not di-rectly related with current and arc voltage andbeing mainly determined by mass feed rate andfractional composition of filler powder as wellas its physical properties. An important problem,considering a fact of application of expensivepowders from nickel- and cobalt-based alloys aswell as special surfacing iron-based alloys as fillermaterials in plasma-powder surfacing, is optimiz-ing of powder consumption and minimizing ofits loss using this method of surfacing [1].

Powder consumption and its loss depend ondesign of some assemblies of surfacing equipment(powder feed, plasmatron), modes of surfacingand, to some extent, structure of part to be sur-faced (geometry of surface). Loss of powder willrise, in particular, if plasmatron is located orperiodically passes close to edge of part beingsurfaced as well as if width of substrate beingsurfaced is less than nozzle diameter.

Figure 1 shows a scheme of drum-type feeder,developed at the E.O. Paton Electric WeldingInstitute [2]. The feeder consists of sealed hous-ing 1, hopper 6 with branch tube, drum 2 andsetting mechanism, made in a form of springloaded bush 4, being moved along the branchtube with the help of lever-screw regulator 3.Mechanism 5 allows for controlling a level ofpowder in the feeder. The drum is actuated byDC motor by means of worm gear (not indicated

in Figure 1). A gap between bush and drum isset in such a way that in the case of fixed drumthe powder will remain in the hopper. Drummovement drags the powder and directs it in areceiving cone, from which powder is transportedthrough tube 7 into plasmatron by transportinggas. Rate of filler powder feeding into plasmatroncan be changed by means of regulation of drumrotation rate (Figure 2). Transporting gas feedsthe filler powder in a gap between focusing 1and plasma-forming 2 nozzles.

Figure 1. Scheme of drum feeder for filler powder feed(designations see in the text)

© E.F. PEREPLYOTCHIKOV, I.A. RYABTSEV, Yu.N. LANKIN, V.F. SEMIKIN and P.P. OSECHKOV, 2014

12/2014 41

The plasmatron has four variants of combina-tion of plasma forming and focusing nozzles dpl ++ df: 2 + 4; 3 + 6; 4 + 7.5; 5 + 9 mm, that alsoallows for regulating powder feed rate. Surfacing

technology [1] determines a variant of optimumcombination of nozzle diameters.

Intensity of surfacing current and consump-tion of transporting gas (Figure 3) are the mainparameters of surfacing mode effecting consump-tion and loss of filler powder. Increase of inten-sity of surfacing current reduces filler powderloss in whole range of rates of its feed (Figure 3,a), since lager amount of powder can be meltedin arc as well as weld pool.

Small consumption of transporting gas (3—4 l/min) often disturbs surfacing process due to

Figure 2. Plasmatron A1756.05 for plasma-powder surfac-ing: 1 – focusing; 2 – plasma-forming nozzle

Figure 4. Cyclogram of process of plasma-powder surfacing:a – surfacing current; b – powder consumption; Ii –current of indirect arc; Im – current of main arc; Id –current at the moment of beginning of drop of powder feedrate; Gc – maximum level of powder consumption (feedrate); ti – duration of increment; ts – duration of soakingin steady process; tc – duration of correction; td – durationof drop

Figure 3. Dependence of coefficient of powder loss ψ onsurfacing current I (a) and consumption of plasma qpl andtransporting qtr gases [1] (b); a – powder feed: – 1.2;

– 2.0; – 3.5; Δ – 6 kg/h; b – 2 kg/hFigure 5. View of A1756 machine for plasma-powder sur-facing with programmer for control of surfacing process

42 12/2014

blocking of plasmatron channels with powder.Loss of powder grows with increase of transport-ing gas consumption (Figure 3, b) as a result ofrising of initial particle rate and deterioration ofconditions of their heating in the arc.

At that, according to the experience, only partof powder particles, moving in a periphery of arccolumn, enters the weld pool. The particles,which fall on surface being surfaced in front ofor from the side of weld pool, are irreversiblylost as a result of rebound resilience from thissurface. Consumption of transporting gas in the

range of 6—9 l/min, at which powder loss doesnot exceed 5—8 %, is supposed to be the optimumone.

There is also loss of powder at the beginningand end of surfacing process, which is difficultto be optimized and controlled, in addition tothe loss, determined by technological peculiari-ties of steady process of plasma-powder surfacing.In theory, a cyclogram of the whole process ofplasma-powder surfacing can be represented inthe following way (Figure 4).

Figure 6. Principal diagram of programmer

Figure 7. Time characteristics of technological parameters of surfacing (a, b) and beads surfaced by these modes (c, d):a, c – current and powder feed were switched-off simultaneously; b, d – powder feed was stopped before currentswitching-off

12/2014 43

Feeding of filler powder should be started af-ter main arc ignition (Figure 4, b). Powder feedis increased during time ti to value Gc determinedby the technology. Crater filling is necessary tobe carried out at the end of surfacing processwith stepwise synchronous reduction of surfacingcurrent and rate of powder feed.

The E.O. Paton Electric Welding Institutedeveloped a programmer for control of plasma-powder surfacing process. It was designed withapplication of microcontroller PIC16F886 (Mi-crochip) and alphanumeric OLED indicatorWEH1601A (Winstar). Microprocessor block ofthe programmer allows for setting the parametersof time characteristics of surfacing current andpowder consumption for plasma-powder surfac-ing machine A1756 (Figure 5) and their optimiz-ing in the process of program control.

Principal diagrams, printed circuit boards,structure and software were developed for theprogrammer. Software was adjusted for techno-logical conditions of A1756 machine:

• technological parameters are set in actualvalues (surfacing current in amperes, powderconsumption in g/min);

• entered are the technological limits for set-ting the parameters of time characteristics 0—250 A and 0—120 g/min.

Scheme of programmer connection to A1756machine with capability of switching of operationfrom programmer as well as in simple design modewas developed and built (Figure 6).

The programmer was technologically testedon A1756 machine by recording the main parame-ters of surfacing mode, i.e. current and powderfeed. Stated problem lied in development of suchmodes of surfacing, which would provide forminimum powder loss and quality formation ofsurfaced beads, including quality crater filling.

If current is switched-off simultaneously orbefore stopping of powder feed, this will resultin bad crater filling and excessive powder loss(Figure 7, a, c).

In the case, when powder feed is stopped be-fore current switching-off, quality cater fillingis provided and powder loss is reduced to theminimum (Figure 7, b, d). Beginning of drop ofcurrent and powder feed are matched in time,however, powder feed is stopped somewhat ear-lier (Figure 7, b).

Thus, developed programmer allows for set-ting the modes, providing good formation of sur-faced beads, in automatic plasma-powder surfac-ing on A1756 machine, quality crater filling andminimum loss of filler powder.

1. Gladky, P.V., Pereplyotchikov, E.F., Ryabtsev, I.A.(2007) Plasma surfacing. Kiev: Ekotekhnologiya.

2. Gladky, P.V., Pereplyotchikov, E.F., Saprykin,Yu.I. et al. (1970) Drum feeder. USSR author’s cert.266111. Int. Cl. 21h 30/17. Otkrytiya. Izobre-teniya, 11, 76.

Received 22.09.2014

44 12/2014

TRAINING OF SPECIALISTS-WELDERS OF KAZAKHSTANIN UKRAINE

One of the tendencies of the modern industrialproduction is the transition to the Internationalnorms and standards. This widens greatly theopportunity of enterprises in cooperation withforeign partners in the scope of the products con-formity to the established requirements. Thesetendencies are especially characteristic of thewelding production, the efficiency of which isdefined significantly by the quality of the per-sonnel professional competence.

The professional training of the personnel, ca-pable to realize the advantages of advanced weld-ing technologies, has its peculiar features, con-nected with the specifics of the welding processand high requirements to the quality of thewelded structures. These features have been re-alized in training programs of the InternationalInstitute of Welding, on the basis of which theinternational system of training and certificationof the personnel is functioning, starting from theworkers-welders and finishing by the diplomaengineers on welding. Due to application of sin-gle training programs and standardized certifica-tion tests, the qualifications in the field of weld-ing are recognized in different countries.

At the present time the training of the per-sonnel of welding production within the framesof the International Certification System is car-ried out by the Authorized National Bodies(ANB), accredited by the IIW, one of which isthe PWI Inter-Industry Training CertificationCenter (IITCC). Since 1998 the Center is anactive participant of the international programsand showed itself as a reliable partner both inUkraine and also abroad. The specialists, trainedby it, are working successfully in many countries.

In April, 2014 a group of trainees of nine citi-zens of Kazakhstan came for training on program«International welder» in accordance with ap-plication of «Transcargo International Ltd.»,participating in the State Program of accelerated

industrial-innovation development of Republicof Kazakhstan. During 16-week training theypassed the full course of professional-theoreticaland practical training, specified by the IIW pro-gram. They mastered the technique of manualarc coated electrode welding of different jointsof flat parts and pipes of structural steels. Fromthe results of the certification tests, includingchecking of theoretical knowledge and practicalhabits, the IITCC examination commissionawarded five trainees the qualification of the In-ternational welder of pipes and four trainees thequalification of the International welder ofplates.

Authority of «Transcargo International Ltd.»recognized the good level of organizing the theo-retical and practical training, high skill of teach-ers and foremen of production training, qualityof methodical provision of training process andexpressed the readiness to continue cooperationin the field of professional training of personnelalso in accordance with other international quali-fications.

Dr. P.P. Protsenko, PWI IITCC

Teachers of IITCC: Drs G.E. Saksonov, E.P. Chvertko andP.P. Protsenko, Director of IITCC (lower row)

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INDEX OF ARTICLES FOR TPWJ’2014, Nos. 1—12Pilot Plant of Welding Equipment of the E.O. Paton Elec-tric Welding Institute at the nowadays stage 4

The 80th Anniversary of the E.O. Paton Electric WeldingInstitute of the NAS of Ukraine 9

VIII INTERNATIONAL CONFERENCEON WELDING CONSUMABLESCONSUMABLES FOR MANUAL ARC WELDINGElectrodes for welding of dissimilar chromium martensiticand chromium-nickel austenitic steels (Zakharov L.S.,Gavrik A.R. and Lipodaev V.N.) 6/7

Heating and melting of electrodes with exothermic mixturein coating (Vlasov A.F., Makarenko N.A. and KushchyA.M.) 6/7

Investigation of composition and structure of weld metalof Kh20N9G2B type made in wet underwater welding(Yushchenko K.A., Bulat A.V., Kakhovsky N.Yu., Samoj-lenko V.I., Maksimov S.Yu. and Grigorenko S.G.) 6/7

Investigation of transition zone of low-carbon steel jointwith high-alloyed Cr-Ni deposited metal (YushchenkoK.A., Kakhovsky Yu.N., Bulat A.V., Morozova R.I., Zvya-gintseva A.V., Samojlenko V.I. and Olejnik Yu.V.) 6/7

Status of normative base, certification and attestation ofwelding consumables in Ukraine (Protsenko N.A.) 6/7

Towards the problem of dispersity and morphology of par-ticles in welding aerosols (Gubenya I.P., Yavdoshchin I.R.,Stepanyuk S.N. and Demetskaya A.V.) 6—7

Ultraviolet radiation in manual arc welding using coveredelectrodes (Levchenko O.G., Malakhov A.T. and ArlamovA.Yu.) 6/7

CONSUMABLES FOR MECHANIZED METHODSOF WELDINGApplication of flux-cored wires for welding in industry(Rosert R.) 6/7

Application of pulse atomizing jet in electric arc metallizing(Royanov V.A. and Bobikov V.I.) 6/7

Development of high-vanadium alloy for plasma-powdersurfacing of knives for cutting of non-metallic materials(Pereplyotchikov E.F.) 6/7

Discrete filler materials for surfacing in current-conductingmould (Kuskov Yu.M.) 6/7

Effectiveness of application of new consumables in weldingand surfacing of copper and its alloys (Review) (IlyushenkoV.M., Anoshin V.A., Majdanchuk T.B. and LukianchenkoE.P.) 6—7

Evaluation of suitability of welding wire of Sv-10GN1MAtype produced by ESAB for manufacturing NPP equipment(Livshits I.M.) 6/7

Flux-cored strips for wear-resistant surfacing (VoronchukA.P.) 6/7

Flux-cored wires for surfacing of steel hot mill rolls (Kon-dratiev I.A. and Ryabtsev I.A.) 6/7

Flux for electric arc surfacing providing high-temperatureremoval of slag coating (Strelenko N.M., Zhdanov L.A.and Goncharov I.A.) 6/7

Influence of active gas content and disperse filler continuityon the process of bead formation in microplasma powdersurfacing of nickel superalloys (Yushchenko K.A. andYarovitsyn A.V.) 6/7

Investigation of influence of microalloying with titaniumand boron of weld metal on its mechanical properties inunderwater welding (Maksimov S.Yu., Machulyak V.V.,Sheremeta A.V. and Goncharenko E.I.) 6/7

Manufacturing defects in welding consumables influencingthe quality of welded joints (Turyk E.V.) 6/7

Materials for strengthening of gas turbine blades (KostinA.M., Butenko A.Yu. and Kvasnitsky V.V.) 6/7

New capabilities of the oldest enterprise on production ofwelding fluxes (Zalevsky A.V., Galinich V.I., GoncharovI.A., Osipov N.Ya., Netyaga V.I. and Kirichenko O.P.) 6/7

Physical-metallurgical and welding-technological proper-ties of gas-shielded flux-cored wires for welding of struc-tural steels (Shlepakov V.N.) 6/7

Restoration and strengthening surfacing of parts of dieequipment (Solomka E.A., Lobanov A.I., Orlov L.N.,Golyakevich A.A. and Khilko A.V.) 6/7

Role of welding flux in formation of weld metal during arcwelding of high-strength low-alloy steels (Golovko V.V.,Stepanyuk S.N. and Ermolenko D.Yu.) 6/7

Selection of shielding gas for mechanized arc welding ofdissimilar steels (Elagin V.P.) 6/7

Tungsten carbide based cladding materials (Zhudra A.P.) 6/7

PROCESSES OF ARC WELDING. METALLURGY.MARKETSApplication of shielding gases in welding production (Re-view) (Paton B.E., Rimsky S.T. and Galinich V.I.) 6/7

Effect of scandium-containing wire on structure and prop-erties of joints of aluminum-lithium alloys produced byargon-arc welding (Markashova L.I., Kushnaryova O.S.and Alekseenko I.I.) 6/7

Effect of structural factors on mechanical properties andcrack resistance of welded joints of metals, alloys and com-posite materials (Markashova L.I., Poznyakov V.D., Berd-nikova E.N., Gajvoronsky A.A. and Alekseenko T.A.) 6/7

Ensuring integrity of welded structures and constructionsat their long-term service with application of renovationtechnologies (Steklov O.I., Antonov A.A. and SevostianovS.P.) 6/7

Interaction of hydrogen with deformed metal (PaltsevichA.P., Sinyuk V.S. and Ignatenko A.V.) 6/7

Investigation of cracking susceptibility of austenitic mate-rial using PVR-test procedure (Yushchenko K.A.,Savchenko V.S., Chervyakov N.O., Zvyagintseva A.V.,Monko G.G. and Pestov V.A.) 6/7

Market of welding consumables in Ukraine (Mazur A.A.,Pustovojt S.V., Petruk V.S. and Brovchenko N.S.) 6/7

Peculiarities of degradation of metal of welded joints ofsteam pipelines of heat power plants (Dmitrik V.V. andBartash S.N.) 6/7

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Underwater welding and cutting in CIS countries(Kononenko V.Ya.) 6/7

TECHNOLOGIES, EQUIPMENT AND CONTROLIN CONSUMABLES PRODUCTIONDirections of improvement of equipment and technologyfor electrode manufacture (Gnatenko M.F., Voroshilo V.S.and Suchok A.D.) 6/7

Effect of charge grain composition on rheological charac-teristics on rheological characteristics of compounds forlow-hydrogen electrodes (Marchenko A.E.) 6/7

Improvement of adaptability to fabrication and weldingproperties of electrodes for tin bronze welding and surfacing(Majdanchuk T.B. and Skorina N.V.) 6/7

State of raw material base of electrode production(Palievskaya E.A. and Sidlin Z.A.) 6/7

Thickness difference of electrode coatings caused by elasticturbulence of electrode compounds under condition of non-isothermal pressure flow (Marchenko A.E.) 6/7

INDUSTRIALAnalysis of some physical and technical characteristics ofion-plasma coating (TiZr)N on rotor blades of compressorof gas-turbine engine TV3-117 (Korsunov K.A. andAshikhmina E.A.) 2

Application of complex-alloyed powders produced by ther-mocentrifugal sputtering in flux-cored wires (Zhudra A.P.,Krivchikov S.Yu. and Dzykovich V.I.) 12

Assessment of effectiveness of residual stress lowering incircumferential joints of pipes after postweld explosiontreatment (Bryzgalin A.G.) 5

Automatic machine for wet underwater welding in confinedspaces (Lebedev V.A., Maksimov S.Yu., Pichak V.G. andZajnulin D.I.) 9

Braze-welded tubular billets for pipelines and high-pressurevessels (Pismenny A.A., Gubatyuk R.S., Prokofiev A.S.,Muzhichenko A.F. and Shinkarenko A.S.) 10

Calculation of upsetting force in flash butt welding ofclosed-shape products (Chvertko P.N., Moltasov A.V. andSamotryasov S.M.) 1

Computer-based technologies and their influence on weld-ing education (Keitel S., Ahrens C. and Moll H.) 10

Development of technologies of repair by arc welding ofoperating main pipelines in Ukraine (But V.S. and OlejnikO.I.) 5

Dry ice – useful material in welding performance (Zhiz-nyakov S.N.) 4

Effectiveness of natural gas transportation by sea at appli-cation of high pressure welded cylinders (Paton B.E., Savi-tsky M.M., Savitsky A.M. and Mazur A.A.) 8

Effectiveness of strengthening butt welded joints after long-term service by high-frequency mechanical peening (KnyshV.V., Solovej S.A. and Kuzmenko A.Z.) 11

Electrode and filler materials for surfacing and welding ofcast tin bronzes (Review) (Majdanchuk T.B.) 1

Electron beam welding of large-size thick-wall structuresof magnesium alloys (Nesterenkov V.M. and BondarevA.A.) 2

Electron beam welding of sheet commercial titanium VT1-0,hardened by nitrogen in the process of arc-slag remelting,and properties of produced joints (Saenko V.Ya., PolishkoA.A., Ryabinin V.A. and Stepanyuk S.N.) 11

Estimation of possibility for producing full-strength jointof large steel parts using the method of autovacuum brazingof threaded profile (Poleshchuk M.A., Atroshenko M.G.,Puzrin A.L. and Shevtsov V.L.) 10

Fatigue calculation for welded joints of bearing elementsof freight car bogie (Lobanov L.M., Makhnenko O.V.,Saprykina G.Yu. and Pustovoj A.D.) 10

Features of reconditioning steel drill bit watercourse (Ste-faniv B.V., Khorunov V.F., Sabadash O.M., MaksymovaS.V. and Voronov V.V.) 11

Flash-butt welding of thin-walled profiles of heat-hardenedaluminium alloys (Chvertko P.N., Semyonov L.A. andGushchin K.V.) 12

Hybrid laser-GMA girth welding technologies for transmis-sion pipelines (Keitel S. and Neubert J.) 4

Industrial electron beam installation L-8 for deposition ofheat-protective coatings on turbine blades (GrechanyukN.I., Kucherenko P.P., Melnik A.G., Kovalchuk D.V. andGrechanyuk I.N.) 10

Joining of thick metal by multipass electroslag welding(Yushchenko K.A., Kozulin S.M., Lychko I.I. and KozulinM.G.) 9

Linear friction welding of metallic materials (Review)(Zyakhor I.V., Zavertanny M.S. and Chernobaj S.V.) 12

Modernization of control system of A1756 machine forplasma-powder surfacing (Pereplyotchikov E.F., RyabtsevI.A., Lankin Yu.N., Semikin V.F. and Osechkov P.P.) 12

Peculiarities of application of supercapacitors in devicesfor pulse welding technologies (Korotynsky A.E., Dra-chenko N.P. and Shapka V.A.) 9

Percussion capacitor-discharge welding of wire of compositesuperconducting alloy (Kaleko D.M.) 4

Producing of bimetal joints by laser welding with full pene-tration (Schmidt M. and Kuryntsev S.V.) 4

Properties of fusion-welded joints on high-strength tita-nium alloy T110 (Akhonin S.V., Belous V.Yu., AntonyukS.L., Petrichenko I.K. and Selin R.V.) 1

Prospects for development of load-carrying elements offreight car bogie (Makhnenko O.V., Saprykina G.Yu., Mir-zov I.V. and Pustovoj A.D.) 3

Reasons of stress corrosion failure of erection girth joint ofmain gas pipeline (Rybakov A.A., Goncharenko L.V., Filip-chuk T.N., Lokhman I.V. and Burak I.Z.) 3

Resistance butt welding of concrete reinforcement in con-struction site (Chvertko P.N., Goronkov N.D., VinogradovN.A., Samotryasov S.M. and Sysoev V.Yu.) 3

Sanitary-hygienic evaluation of noise in manual arc weldingwith covered electrodes (Levchenko O.G., Kuleshov V.A.and Arlamov A.Yu.) 9

State-of-the-art and prospects of world and regional marketsof welding materials (Review) (Mazur A.A., PustovojtS.V., Makovetskaya O.K., Brovchenko N.S. and PetrukV.S.) 11

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Structure and properties of welded joints of 15Kh1M1FLsteel at repair of casting defects by transverse hill method(Efimenko N.G., Atozhenko O.Yu., Vavilov A.V., KantorA.G. and Udalova E.I.) 2

Technological peculiarities of laser microplasma and hybridlaser-microplasma welding of aluminium alloys (ShelyaginV.D., Orishich A.M., Khaskin V.Yu., Malikov A.G. andChajka A.A.) 5

Technology for manufacture of gas-and-oil line pipes usinghigh-frequency method of welding at Company «InterpipeNMPP» (Antipov Yu.N., Dmitrenko E.V., KovalenkoA.V., Goryanoj S.A., Rybakov A.A., Semyonov S.E. andFilipchuk T.N.) 3

Wear-resistant arc surfacing over the layer of alloyingcharge (Peremitko V.V.) 8

Welded structure of Kiev TV-tower is 40 years old (Lo-banov L.M., Garf E.F., Kopylov L.N. and Sineok A.G.) 1

55th ANNIVERSARY OF WELDINGPRODUCTION CHAIR OF ADMIRAL MAKAROVNATIONAL SHIPBUILDING UNIVERSITYEffect of stress-strain state on structure and properties ofjoints in diffusion welding of dissimilar metals (KvasnitskyV.V., Kvasnitsky V.F., Markashova L.I. and MatvienkoM.V.) 8

Effect of weld convexity sizes on stress state of butt jointduring tension (Ermolaev G.V., Martynenko V.A. andMarunich I.V.) 8

Increase of service properties of electric-arc and plasmacoatings by use of electric-pulse effect on double-phase high-temperature flow (Dubovoj A.N., Karpechenko A.A. andBobrov M.N.) 8

Instability of mode in circuit with capacity and electric arcsupplied by direct current source (Vereshchago E.N. andKostyuchenko V.I.) 8

Regularities of creation of modified interlayers in using ofhighly-concentrated energy flows (Kvasnitsky V.F., Kvas-nitsky V.V., Cherenda N.N., Koval N.N. and LevchenkoI.L.) 8

Stress-strain state at force and temperature loading of as-sembles from dissimilar steels with soft interlayer (KolesarI.A. and Ermolaev G.V.) 8

Technological characteristics of automatic submerged arcsurfacing with high-frequency oscillations of electrode end(Lebedev V.A., Dragan S.V., Goloborodko Zh.G.,Simutenkov I.V. and Yaros Yu.A.) 8

INFORMATIONPaton Turbine Technologies – new name of a well-knowncompany 9

The 55th Anniversary of the Experimental Design Techno-logical Bureau of the E.O. Paton Electric Welding Institute 9

Today’s Foreign Trade Company «INPAT» of the E.O.Paton Electric Welding Institute 2

Training of specialists-welders of Kazakhstan in Ukraine 12

20 years in the world of flux-cored wires 1

NEWSInternational Conference «Welding Consumables» 8

International Conference «Welding and Related Technolo-gies – Present and Future» 1

The Evgeny Paton Prize Winners of 2013 5

SCIENTIFIC AND TECHNICALAnalysis and procedure of calculation of series connectionelectronic devices for contactless arc excitation (MakhlinN.M. and Korotynsky A.E.) 1

Cermet coatings of chromium carbide-nichrome system pro-duced by supersonic plasma gas air spraying (Korzhik V.N.,Borisova A.L., Popov V.V., Kolomytsev M.V., ChajkaA.A., Tkachuk V.I. and Vigilyanskaya N.V.) 12

Cracking susceptibility of welded joints in repair structureson main gas pipelines (But V.S., Maksimov S.Yu. andOlejnik O.I.) 11

Deformation-free welding of stringer panels of titaniumalloy VT20 (Paton B.E., Lobanov L.M., Lysak V.L., KnyshV.V., Pavlovsky V.I., Prilutsky V.P., Timoshenko A.N.,Goncharov P.V. and Guan Qiao) 9

Determination of the causes for crack initiation in structuralelements of the tower of new ventilation pipe at ChernobylNPP (Torop V.M., Garf E.F., Yakimkin A.V. and GopkaloE.E.) 1

Detonation coatings of composite powder of ferromolybde-num-silicon carbide produced using method of mechanical-and-chemical synthesis (Borisov Yu.S., Borisova A.L., As-takhov E.A., Burlachenko A.N., Ipatova Z.G. and GorbanV.F.) 3

Effect of cyclic load on microstructure and cold resistanceof the 10G2FB steel HAZ metal (Poznyakov V.D.,Markashova L.I., Maksimenko A.A., Berdnikova E.N.,Alekseenko T.A. and Kasatkin S.B.) 5

Effect of electric parameters of arc surfacing using flux-cored wire on process stability and base metal penetration(Lankin Yu.N., Ryabtsev I.A., Soloviov V.G., ChernyakYa.P. and Zhdanov V.A.) 9

Effect of nickel and manganese on structure of Ag—Cu—Zn—Sn system alloys and strength of brazed joints (KhorunovV.F., Stefaniv B.V. and Maksymova S.V.) 4

Effect of temperature of thermomechanical treatment onquality of dissimilar metal joints (Demidenko L.Yu., Ona-tskaya N.A. and Polovinka V.D.) 12

Elimination of local deformations of buckling type by meansof electrodynamic treatment (Lobanov L.M., PashchinN.A., Mikhoduj O.L. and Solomijchuk T.G.) 11

Fatigue life of deposited repair welds on single-crystal high-temperature nickel alloy under cyclic oxidation (BelyavinA.F., Kurenkova V.V. and Fedotov D.A.) 2

Features of melting, structure and properties of Ni—Mn—Cusystem nickel alloys (Khorunov V.F. and Lototsky P.N.) 5

High-power laser welding of austenitic stainless steel withelectromagnetic control of weld pool (Bachmann M.,Avilov V., Gumenyuk A. and Rethmeier M.) 3

Improvement of power efficiency of machines for resistancespot welding by longitudinal compensation of reactivepower (Pismenny A.A.) 1

Improvement of the procedure of mode parameter calcula-tion for gas-shielded multipass welding (Buzorina D.S.,Sholokhov M.A. and Shalimov M.P.) 10

Increase of fatigue resistance of sheet welded joints of alu-minum alloys using high-frequency peening (Knysh V.V.,Klochkov I.N., Pashulya M.P. and Motrunich S.I.) 5

Influence of heating rate on inflammation temperature ofmutlilayer Ti/Al foil (Kuzmenko D.N., Ustinov A.I., Ko-sintsev S.G. and Petrushinets L.V.) 10

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Influence of non-metallic inclusions in pipe steels ofstrength class X65—X80 on values of impact toughness offlash-butt welded joints (Kuchuk-Yatsenko S.I., ShvetsYu.V. and Shvets V.I.) 12

Influence of non-uniformity of heating on upsetting forcevalue and forging time in flash-butt welding of flat ring(Moltasov A.V., Samotryasov S.M., Knysh V.V., ChvertkoP.N. and Gushchin K.V.) 10

Integrated evaluation of effect of main impurities on weld-ability of copper (Anoshin V.A., Ilyushenko V.M., Bon-darenko A.N., Lukianchenko E.P. and Nikolaev A.K.) 11

Investigation of spraying spot and metallization patternunder conditions of microplasma spraying of coatings oftitanium dioxide (Borisov Yu.S., Vojnarovich S.G., Kis-litsa A.N. and Kalyuzhny S.N.) 12

Investigation of thermal resistance of deposited metal de-signed for restoration of mill rolls (Babinets A.A., RyabtsevI.A., Kondratiev I.A., Ryabtsev I.I. and Gordan G.N.) 5

Investigation of wear resistance of composite alloys underthe conditions of gas-abrasive wear at elevated temperatures(Zhudra A.P.) 11

Laser and laser-microplasma alloying of surface of38KhN3MFA steel specimens (Shelyagin V.D.,Markashova L.I., Khaskin V.Yu., Bernatsky A.V. andKushnaryova O.S.) 2

Methods of mathematical modelling of the processes of elec-trode metal drop formation and transfer in consumable elec-trode welding (Review) (Semyonov A.P.) 10

Microstructure of brazed joints of nickel aluminide (Mak-symova S.V., Khorunov V.F., Myasoed V.V., VoronovV.V. and Kovalchuk P.V.) 10

Modeling of weld pool behaviour in spot welding by pulsedlaser radiation (Semyonov A.P., Shuba I.V., Krivtsun I.V.and Demchenko V.F.) 4

Numerical modeling and prediction of weld microstructurein high-strength steel welding (Review) (Ermolenko D.Yu.and Golovko V.V.) 3

Peculiarities of alloying of weld metal of high-strengthaluminium alloy welded joints with scandium (FedorchukV.E., Kushnaryova O.S., Alekseenko T.A. and FalchenkoYu.V.) 5

Peculiarities of structure of coatings of Fe—Cr—Al systemflux-cored wire produced under conditions of supersonicelectric arc metallization (Korzhik V.N., Borisova A.L.,Gordan G.N., Lyutik N.P., Chajka A.A. and Kajda T.V.) 2

Prediction of thermodynamical properties of melts of CaO—Al2O3 system (Goncharov I.A., Galinich V.I., MishchenkoD.D. and Sudavtsova V.S.) 4

Quasi-crystalline alloys-fillers for composite layers pro-duced using method of furnace surfacing (Sukhovaya E.V.) 1

Redistribution of residual welding stresses in in-vessel corebarrel of WWER-1000 reactor during operation (Makh-nenko O.V., Velikoivanenko E.A. and Mirzov I.V.) 11

Resistance to cold crack formation of HAZ metal of weldedjoint on high-strength carbon steels (Gajvoronsky A.A.) 2

Simulation of electric arc with refractory cathode andevaporating anode (Krikent I.V., Krivtsun I.V. and Dem-chenko V.F.) 9

Structural changes in overheating zone of HAZ metal ofrailway wheels in arc surfacing (Gajvoronsky A.A., ZhukovV.V., Vasiliev V.G., Zuber T.A. and Shishkevich A.S.) 1

Structural features of FSW joints of metals with differentelement solubility in the solid phase (Grigorenko G.M.,Adeeva L.I., Tunik A.Yu., Stepanyuk S.N., PoleshchukM.A. and Zelenin E.V.) 4

Structure of surface-melted zone of cast high-nickel alloyKhN56MBYuDSh after laser surface treatment (PolishkoA.A., Saenko V.Ya., Tunik A.Yu. and Stepanyuk S.N.) 3

System for measurement of temperature of biological tissuesin bipolar high-frequency welding (Lankin Yu.N., SushyL.F. and Bajshtruk E.N.) 11

Technological peculiarities of welding of wrought magne-sium alloys by electron beam in vacuum (Bondarev A.A.and Nesterenkov V.M.) 3

Tribotechnical properties of deposited metal of 50Kh9S3Gtype with increased sulphur content (Osin V.V.) 12

Index of articles for TPWJ’2014, Nos. 1—12 12

List of authors 12

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LIST OF AUTHORS

Adeeva L.I. No.4

Ahrens C. No.10

Akhonin S.V. No.1

Alekseenko I.I. Nos. 6/7

Alekseenko T.A. No.5(2), 6/7

Anoshin V.A. No.6/7, 11

Antipov Yu.N. No.3

Antonov A.A. Nos. 6/7

Antonyuk S.L. No.1

Arlamov A.Yu. No.6/7, 9

Ashikhmina E.A. No.2

Astakhov E.A. No.3

Atozhenko O.Yu. No.2

Atroshenko M.G. No.10

Avilov V. No.3

Babinets A.A. No.5

Bachmann M. No.3

Bajshtruk E.N. No.11

Bartash S.N. Nos. 6/7

Belous V.Yu. No.1

Belyavin A.F. No.2

Berdnikova E.N. No.5, 6/7

Bernatsky A.V. No.2

Bobikov V.I. Nos. 6/7

Bobrov M.N. No.8

Bondarenko A.N. No.11

Bondarev A.A. No.2, 3

Borisov Yu.S. No.3, 12

Borisova A.L. No.2, 3, 12

Brovchenko N.S. No.6/7, 11

Bryzgalin A.G. No.5

Bugaenko B.V. No.8

Bulat A.V. Nos. 6/7(2)

Burak I.Z. No.3

Burlachenko A.N. No.3

But V.S. No.5, 11

Butenko A.Yu. Nos. 6/7

Buzorina D.S. No.10

Chajka A.A. No.2, 5, 12

Cherenda N.N. No.8

Chernobaj S.V. No.12

Chernyak Ya.P. No.9

Chervyakov N.O. Nos. 6/7

Chvertko P.N. No.1, 3, 10, 12

Demchenko V.F. No.4, 9

Demetskaya A.V. Nos. 6/7

Demidenko L.Yu. No.12

Dmitrenko E.V. No.3

Dmitrik V.V. Nos. 6/7

Drachenko N.P. No.9

Dragan S.V. No.8

Dubovoj A.N. No.8

Dzykovich V.I. No.12

Efimenko N.G. No.2

Elagin V.P. Nos. 6/7

Emelianov V.M. No.8

Ermolaev G.V. No.8(2)

Ermolenko D.Yu. No.3, 6/7

Falchenko Yu.V. No.5

Fedorchuk V.E. No.5

Fedotov D.A. No.2

Filipchuk T.N. No.3(2)

Gajvoronsky A.A. No.1, 2, 6/7

Galinich V.I. No.4, 6/7(2)

Garf E.F. No.1(2)

Gavrik A.R. Nos. 6/7

Gnatenko M.F. Nos. 6/7

Goloborodko Zh.G. No.8

Golovko V.V. No.3, 6/7

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